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SPECIFICATION AND DESIGN

DYNAMO-ELECTRIC MACHINERY

LONOMA N8'
ELECTRICAL ENGINEERING SERIES.

Edited by CHARLES P. SPARKS. M.Inst.C.E., M.I.E.E.

POWER HOUSE DESIGN. By John F. C. Snell,
M.Inst.C.E., M.I.E.E. With 17 Folding Plates and
186 Illustrations. 21s. net.

SPECIFICATION AND DESIGN OF DYNAMO-
ELECTRIC MACHINERY. By Miles Walker,
M.A., M.I.E.E. 32s. net.

THE DIAGNOSING AND CURING OF TROUBLES
IN ELECTRICAL MACHINES. By Miles
Walker, M.A., M.I.E.E. [/n preparation,

THE APPLICATIONS OF ELECTRICITY TO

FACTORIES. By John Shaw. M.I.E.E.

[/» preparation,

ELECTRICAL EQUIPMENT OF MINES. By Charles
P. Sparks, M.Inst.C.E., M.I.E.E. \In preparation.

LONGMANS, GREEN & CO.
London, New York, Bombay, CAixurrA and Madras.

SPECIFICATION AND DESIGN

DYNAMO-ELECTRIC
MACHINERY

MILES WALKER, M.A., M.I.E.E.

PROFESSOR OF ELBOTRICAL ENQINKBRING IN THK FACULTT OF TECHNOLOOT IN THE
UNIYERSITT OF MANCHESTER (MANCHESTER SCHOOL OF TECHNOLOGY)

CONSULTING DESIGNER TO THE BRITISH WESTINGHOUSE ELECTRIC AND MANUFACTURING CO. LTD.

l^ITH ILLUSTRATIONS

LONGMANS, GREEN AND CO.

39 PATERNOSTER ROW, LONDON

FOURTH AVENUE k SOth STREET, NEW YORK

BOMBAY, CALCUTTA, AND MADRAS

1915

200201 a -7/0 w

DEC 31 19:5 lO-i [ ^blS

TO ^

■VSJ\5

PREFACE

The books on the design of dynamos are so numerous and so excellent, that a
serious apology is necessary for adding another to our crowded shelves. When
the author was asked to write a book on Design for Messrs. Longmans' Electrical
Engineering Series, he was in doubt whether he should take up the task. There
appeared, however, to be no book of precedents of electrical specifications
analogous to the famous "Conveyancing Precedents" compiled by Prideaux^
which are so widely used by lawyers; and it occurred to the author that such
a book would be of some use to those engineers who have from time to time
to draw up specifications for the purchase of electrical machinery.

But a book of precedents alone would be incomplete unless it showed how
the specifications might be fulfilled in the factory; and the author therefore
proposed to add to each specification a worked-out design, showing at least one
method of meeting the prescribed conditions. In order to do this, it has been
necessary to give in the first part of the book a collection of simple rules for
calculating the dimensions and quantities met with in dynamo-electrical machinery.
The general method of design is that employed by many of the engineers of the
Westinghouse Companies of America and Great Britain, who learnt it from
Mr. B. G. Lamme. The advantages of the method are set out on pages 7 and 8.
Many additions and refinements have been made by various users, so that the
rules given are very much more complicated than in the original scheme; but
the beauty of the method is that these refinements can be used or not, according
as the time available for the work is long or short. Most commonly a calculation
sheet, instead of being filled up like those given in the text, contains only a
dozen figures or so, which represent the design sufficiently well for the purpose
of quoting a price.

It would take many years to compile a satisfactory book of precedents; for
it is only by actual experience with the requirements of machinery intended to
work under the many conditions met with in practice, that one can foresee all
the qualities that should be asked for from the maker. For this reason, as a

vi PREFACE

first attempt the author has confined himself to some of the more usual types
of machines, and has left for future consideration specifications relating to the
more special machinery required in mines, rolling mills and factories.

As the book has already exceeded considerably the size originally planned,
a great deal of information which is commonly found in books on design has
been intentionally omitted; and only those tables are included which contain
information not so easily accessible elsewhere.

The author is indebted to Mr. V. M. Allen for most of the matter contained in
Chapter VIL on the design of armature coils and formers ; to Mr. K. Faye-Hansen
for Figure 312; to Mr. S. C. Nottage for Figure 405; to Dr. W. Petersen for
permission to use Figures 217, 220, 221, 302, 306, 307 to 309, 311, 315, 316,
348, 349, 354 to 358, 361, 376 to 378, 402, 404 and 411 from his book
Wechselstrommaschinen ; and to Dr. E. Rosenberg, Mr. J. W. Schrooder, and
Mr. Robert Townend for valuable criticisms and suggestions. He also wishes
to thank Mr. David Isaacs for his indefatigable proof-reading and the preparation
of the Indexes. A great number of the drawings have been specially made for
the book by Mr. J. Mitscha. Lastly, the author wishes to express his thanks
to the British Westinghouse Electric & Manufacturing Company, Limited; for
their permission to publish some of the information contained in the book.

Mancuesteb, June 1915.

CONTENTS

PAOK

Tables ----... xii

Specifications and Calculation Sheets xiii

Symbols xv

PART I. SHORT RULES FOR USE IN THE DESIGN
OF DYNAMO'ELECTRIC MACHINERY.

CHAPTER

I. Introduction.

One general method for all machines 4

Fmidamental formula for voltage generation 5

Electromotive force coefficient Ke 7

II. The Magnetic Circuit.

Effect of number of poles on general design ICT

The field form, and field-form coefficient Kf 13

of salient poles 13.

of distributed winding 18-

of induction motor ZO*

III. The Magnetic Circuit (continued).

Calculation of the E.M.F. coefficient Ke 23:

for a C.C. machine 23-

for an A.C. machine 24

for an induction motor . - - . • - 3^

IV. The Materials of the Magnetic Circuit.

Magnetic units

Magnetic properties of iron and steel -
Losses in sheet iron - - - -

Hysteresis losses

Eddy-current losses -.--'-

35
36
45
45

viii CONTENTS

CHAFTKR PACK

V. The Parts of the Magnetic Circuit.

The air-gap 56

Slots and teeth .-- 67

Air-gap-and-tooth-saturation curve 76

Design of teeth - 79

The iron behind the slots 82

The yoke 85

VI. The Electric Circuits.

Armature windings 87

The conductor diagram and* winding diagram 87

Classes of 3-phase armature windings 101

The effect of chording the winding - - Ill

Mechanical arrangement of windings - • -. - - - - 115

The current in short-circuited windings • 123

Switching in alternators when out of step 132

Material of conductors 135

Shape of conductors 138

Size of conductors 141

Eddy-currents in armature conductore - 144

VII. The Design of Armature Coils and the Formers on

WHICH THEY ARE WOUND.

Lattice coils 151

Pulled coils 155

Diamond type coils 156

Short type coils 163

Concentric coils - - 168

Field moulds 172

VIII. Insulation.

Mechanical qualities 175

Dielectric strength 186

Pressure tests - - 187

Specific resistance and effect of moisture 189

Withstanding high temperatures 190

Heat conductivity 191

Oxidization and slow changes with time 191

Formation of nitric acid - 192

Experience from breakdowns 193

Method of insulating coils 197

Room taken up by insulation - - - 201

IX. Ventilation.

Effect of general shape of the franic -
Amount of air required
Schemes of ventilation

204
206
206

CONTENTS ix

CHAPTER PAOE

Radial ducts and axial ducts 208

Power taken to drive fan 213

Friction and windage losses 216

* X. The Predetermination of Temperature Rise.

Conduction of heat 219

Cooling by air - ^ - 229

Conduction of heat across the layers of a coil 236

Passage of heat from the surface of ventilating ducts . . - - 241

Conductivity of iron punchings - 251

Cooling of external surface 254

Collection of rules for predetermining the cooling conditions - - 254

Permissible temperatures 256

PART IL THE SPECIFICATION AND THE DESIGN
TO MEET THE SPECIFICATION.

XI. The Specification and the Design to meet the Speci-
fication.

Performance specifications in general 261

Arrangement of clauses 262

XII. Alternating-Current Generators — High-Speed Engine
Type.

Specification No. 1, 750-K.V.A. 3- phase engine-driven generator 269

The design to meet the specification 274

The regulation of A.C. generators - - - - 278

Arrangement of copper and iron ..... 300

The wave form of the electron^i^tive force 304

Calculation of a 750-K.V.A. engine driven generator - - - 316

Calculation sheet No. 1 for this machine 321

XIII. Alternating-Current Generators (continued) ^SIjOW-

Speed Engine Type.

Specification No. 2, 2180 K.V.A. 3-pha8e generator direct-connected

to gas engine 333

The design of this machine 337

Parallel running of alternators 337

Method of fixing on size of flywheel for alternator driven by a prime

mover of irregular turning moment 344

Design of 2180-K.V. A. generator to meet Specification No. 2 - - 347

Calculation sheet No. 2 for this machine 348

Effect of higher speed on the design 356

Calculation sheet No. 3, 1800-K.V.A. 3-phase A.C. generator, 150

R.P.M. 357

X CONTENTS

CHAPTSR PAOB

XIV. Alternating-Current Generators (continued) — Water-

Turbine Type.

Specification No. 4, 2500-K.V.A. 3-pha49e generator to be driven by

water turbine 369

Design of machine to meet Specification No. 4 361

Calculation sheet No. 4 for this machine 364

XV. Alternating-Current Turbo-Generators.

Centrifugal forces on rotor 367

Different types of field-magnet - - - - . - - - 368

Rotor windings 371

Field-form of cylindrical field-magnet 375

Specification No. 6, 15,000-K.V.A. 3-phase turbo-generator - - 378

Design of this machine 383

Calculation sheet No. 6 for this machine ... . - 387

Two-pole turbo-generators 402

Specification No. 6, 2600-K.V.A. 3-phaae turbo-generator - - - 404

Design of this machine 405

Calculation sheet No. 6 for this machine - - • - - 408

25-cycle turbo-generators 409

Single- phase generators 411

XVI. Induction Motors.

The circle diagram 413

Determination of the magnetizing current of an induction motor - 416
Determination of the short-circuit current by calculation from the

design 420

The reactance of the motor on short circuit 428

The apparent resistance of the motor on short circuit - - 428

The apparent impedance of the motor on short circuit - • - 428

Power factor for various values of r and various loads - - - 429

Crawling of induction motors 429

Slip of induction motors 433

XVII. Induction Motors {continued.)

Specification No. 7, 1500-H.P. 3-phase induction motor, 246 R.P.M. - 438

The design of this machine 445

Calculation sheet No. 7 for this machine 448

Specification No. 8, 350-H.P. induction motor, 1480 R.P.M. - - 460

The design of this machine 462

Calculation sheet No. 8 for this machine 463

Small motors 467

Specification No. 9, 36-H.P. induction motor, 960 R.P.M. - - - 468

The design of this machine 470

Calculation sheet No. 9 for this machine 471

CONTENTS xi

CHAPTER PAGE

XVIII. CONTINUOUS-CUREBNT GENERATORS.

The specification of G.C. generators 484

Specification No. 10, 75-K.W. beltdriven C.C. generator, 750 R.P.M. - 486

The design of this machine 487

Calculation sheet No. 10 for this machine 489

Specification No. 11, 1000-K.W. C.C. generator to form part of a motor-
generator set, 246 R.P.M. 500

Calculation sheet No. 11 for this machine 504

The design of this machine 505

Special C.C. generators - - - 510

Arnold singly re-entrant multiplex winding 511

Calculation sheet No. 12, 200-K.W. C.C. generator - - - - 514

The specification of C.C. turho-generators 516

Specification No. 13, steam-turbine C.C. generator set - - . 519

The design of a 1000-K.W. C.C. turbo-generator .... 529

Calculation slieet No. 13 for this machine 531

XIX. Rotary Converters.

Specification No. 14, 1250-K.W. rotary converter and A.C. booster - 560

The design of this converter 567

Calculation sheet No. 14 for this converter 570

The design of an A.C. booster 579

Calculation sheet No. 14a, 115-K.V.A. 6-phase A.C. booster - - 582

Large low- voltage converters 583

Specification No. 15, 2000-K.W. rotary converter, 250 volts - - 584

The design of a 2000-K.W. rotary converter for electrolytic work - 594
The variation of the voltage of a rotary converter by the variation of

its excitation 595

Special precautions necessitated when the frequency is unsteady - 600
The phase-swinging of synchronous motors and rotary converters,

with and without dampers 602

Small rotary converters 604

XX. Phase Advancers.

Specification No. 7a, 1500-H.P. 3-phase induction motor intended to

be run on leading power factor 608

Specification No. 16, 30 K.V.A. phase advancer 610

The design of this machine 612

Calculation sheet No. 16 for this machine 617

Change of speed of induction motors 623

Index OF the Clauses in the Specifications- - - 627

General Index 634

TABLES

TABLK PArsK

I. Hysteretic Constants 48

II. Winding Table : Wave winding, Glass B 104

III. Winding Table : Wave winding, Class C 105

IV. Winding Table : Wave winding, Class C^ 106

V. Winding Table : Wave winding, Class D 107

VI. Winding Table : Wave winding, Class E 107

VII. Giving numbers of Slots that can be used with a given number of Poles

to form a symmetrical 3- phase Winding, two CVinductors per Slot - 109

VIII. Dimensions of the overhang of Concentric Coils 172

IX. The Qualities of Insulating Materials 176

X. Allowance of room in Slot for the external Wrapping of Armature Coils

of A.C. Generators and Motors 202

XI. Power taken to drive Fans 213

XII. Heat Conductivity of Metals 220

XIII. Heat Conductivity of Insulating Materials 221

XIV. Value of hu for Wire- wound Coils 239

XV. Values of Winding Factors, Uniformly Distributed Windings - - 307

XVI. Winding Factors for Phase E.M.F. of 3-pha8e Windings in Slots - - 313

XVII. Values of K^ for different Ratios of Pole-arc to Pole-pitch - - - 342

XVIII. Values oi Ki for End Leakage of 3-phase Motors with normal Full-pitch

Windings 427

XIX. Ratings of Frames of 50-cycle, S-phase Induction Motors - - - 447

XX. Winding Table of 200.K.W. Generator with Arnold Multiplex Singly

Re-entrant Winding 515

XXI. Showing arrangement of Equalizing Connections of Arnold Multiplex

Singly Re-entrant Winding .- 515

XXII. Ratios of C.C. to A.C. Voltage on Rotary Converters as affected by

the Ratio of Pole-arc to Pole-pitch 540

XXIII. Ratios of A.C. Amperes per Slip-ring to C.C. Amperes per Terminal,

assuming an Efficiency of 95 per cent; 541

SPECIFICATIONS AND CALCULATION SHEETS

Speci- Calculation
No. Detaii^. fication Sheet

Page Pago

1. 750-K.V.A. 3-phase engine-driven generator, 375 R.P.M.,

2000-2100 volts, 50 cycles 269 321

2. 2180-K.V.A. 3-phase generator, 6300 volts, 50 cycles, 125

R.P.M., to be direct-connected to a gas engine - - 333 348

3. See Specification No. 2 and page 356, 1800-K.V.A. 3-phase

generator, 63006600 volts, 50 cycles, 150 R.P.M. - — 357

4. 2500-K.V.A. 3-phase generator, 6900 volts, 50 cycles, to be

driven by a water- turbine at 600 R.P.M. - - 359 364

5. 15,000-K.V.A. 3-phase turbogenerator, 11,000 volts, 60

cycles, 1500 R.P.M. 378 387

6. 2500-K.V.A. 3-pha8e turbo-generator, 550 volte, 50 cycles,

3000R.P.M. 404 408

7. 1500-H.P. 3-pha8e 1350-K.V.A. induction motor, 3000 volte,

50 cycles, 246 R.P.M. 438 448

la, 1500-H.P. 3-pha8e induction motor intended to be run on

leading power-factor 608 448

8. 350-H.P. 305-K.V.A. 3-phase induction motor, 2200 volte,

50 cycles, 1480 R.P.M., for pump driving - - - 460 463

9. 35-H.P. 34-K.V.A. 3-phase induction motor, 500 volte, 50

cycles, 980 R.P.M. 468 471

10. 75-K.W. belt-driven C.C. generator, 525 volts, 25 cycles,

750R.P.M. 486 489

11. 1000-K.W. C.C. generator, 460-500 volts, 25 cycles, 246

R.P.M., to form part of a motor-generator set - - 500 504

12. See Specification 11 and page 510, 200-K.W. C.C. generator,

250 volts, 12 cycles, 180 R.P.M. — 514

13. 1000-K.W. C.C. turbo-generator, 550-600 volte, 92 cycles,

2750R.P.M. 519 631

xiv SPECIFICATIONS AND CALCULATION SHEETS

Speci- Calculation
No. Dktails. ^ fication Sheet

* Page Page

14. 1260-K.W. 6-phase rotary converter, 460-560 volts, 50

cycles, 428 RP.M., and A.C. booster - - - - 560 570

14a. 115-K.V.A. 6-phase A.C. booster, 28 volts, 60 cycles, 428

R.P.M. 660 582

16. 2000-K.W. 6-pha8e rotary converter, 260 volts, 60 cycles,

250 R.P.M. 584 —

16. 30-K.V.A. 3-phase advancer, 60 volts (70 max.), 0'66

cycle, 760 R.P.M. 610 617

SYMBOLS

SYMBOL PAOK

Ag = area of working face in cms. =27rrZ 5

Ap = „ „ in inches 6

AgB = total maximum flux of whole frame 6

a = 2mr«-rd81 356 and 601

2a = number of armature cirouite in parallel 512

at = average overhang of coils in cms. 388

B = magnetic flux-density in CG.S. lines per sq. cm. 5

B' = magnetic flux-density in lines per sq. inch 6

Be = B in gap necessary for good commutation 480

Bg = flux-density in the air-gap 308

Ba = coefficient of the A^^ harmonic in the expansion of Bj - - - - 311

Bjc = Kapp lines per sq. inch 6

Bmax = maximum magnetic flux-density per sq. cm. 49

b = width of slot 79

bp = breadth of brush increased in ratio da-rdc 479

c = distance from comer of pole to neutral line 14

e = number of paths in parallel 24

c = drop of core below bore of iron 162

Cp = width of commutator bar increased in ratio dg-^dc 479

D — diameter of armature in cms. 154

ly = diameter of armature in inches 299

Z),n = diameter of armature in metres 479

Dr = greatest deflection of rotor in inches 405

d — depth of winding in cms. 239

de = diameter of commutator 479

E = electromotive force in volts 4

E = voltage of network 339

Ea = voltage to star-point 429

Et = terminal voltage 342

e = instantaneous E.M.F. generated 306

xvi SYMBOLS

SYMBOL PAOK

eg = evanescent voltage - • - - 128

eg = voltage after continued short circuit 1 28

/ = depth of conductor in cms. 147

/q = frequency of oscillation 346

0 = kilograms mass of flywheel 341

g = length of air-gap between pole^and armature 58

H = intensity of field 36

H^ = s^h^k* 543

h = height of slot 79

h = number of the h^^ harmonic 307

h = ratio of A.C. power to C.C. power in rotary converter - . . , 543

he = height of conductors 79

hfi = cooling coefficient, watts per sq. cm. per degree difference of temperature

passing from surfeu^e cooled by a draught of air 229

he = cooling coefficient for ends of coils 233

hi = cooling coefficient for sides of coil 233

hf, = cooling coefficient for ventilating ducts 211

hy = cooling coefficient for cylindrical surface of armature .... 229

hp = total height of pole 238

/ = electric current in amperes - - 218

I A = armature current 7

la = current per conductor 8

I^Za = current loading 8

Id = current density in amperes per sq. cm. - - 227

Ig = magnetizing eddy-current 128

// = field current 299

Ijc = power factor = power factor on short circuit 299

Ii = full-load current 341

Ifn = magnetizing current 420

In = no-load current 429

Inl = current per phase supplying no-load losses 420

Iq = short-circuit current 342

I^fc = short-circuit current 421

Igf^ = instantaneous current flowing when generator is short circuited - - 131

It = termina amperes 512

/{( = current flowing in alternator for unit displacement of the pole centre • 339

Ix see page 239

in = thickness of insulation per cm. of depth of winding 239

Ka = Carter's coefficient for flux fringing from poles 17

Kd = Field's coefficient = ratio of actual loss in conductor to loss there would

be if no eddy-current 146

Kf = electromotive force coefficient 5

SYMBOLS xvli

sy>:bol paok

Kf = flux coefficient 16

Kg = air-gap coefficient 65

iT^ = hysteresis coefficient - 47

K}t = heat conductivity of iron punching in calories per second per sq. cm. per

**C. percm. 253

Kl — leakage coefficient for end- windings 388

K^ = number of commutator bars 512

K^ = output coefficient 447

Kq = internal displacement coefficient 294

Kt = regulation coefficient 299

Kf = space coefficient = ratio of iron + air space to iron space - - - - 71

Kt = zigzag leakage coefficient 424

K^ — cross-magnetizing coefficient 342

h — ratio of wattless current to power current at unity efficiency - - - 643

hh — heat conductivity of insulation in watts per sq. cm. per ° C. per cm. of path 239

L — inductance in henries 129

Le = flux leaking across to the commutating pole and back again - - 480

Xjfc = flux leaking from top of teeth along air-gap 480

Z»„ = effective flux crossing body of slot 480

Ln ' = flux encircling end-connections of armature coil 480

Li — sum of leakage fluxes per cm. of iron . - - .... 480

I = axial length of working face in cms. 5

I = length of bobbin in cms. 239

Zj = coefficient of self-induction 129

/fl = length of path through armature core 55

If = effective axial length of pole 328

Ip — pitch of poles 426

It = length of turn 161

ly = length of yoke 55

Ig = length of teeth 55

Jf = magnetomotive force in C.G.S. units 36

Mp = magnetic potential 5ft

m = width of mouth of slot 79

m — number of slip-rings of rotary converter 643

iV = total magnetic flux per pole 4

N^ = number of slots in armature 612

n = revolutions of armature per second 5

n = frequency in cycles per second 49

n^ = frequency of disturbance 339

n^ — frequency of phase-swing 33ft

p = smallest pitch of slot on cylindrical surface 15ft

/?, = coefficient controlled by heat gradient 23ft

xvui •

SYMBOLS

R =

Rna =

►JW

r
r

SYMBOL PAHK

= number of poles 299

= pitch of poles 328

= pitch of slots C6

= slots per pole 309

= disturbing torque 339

= synchronizing torque 339

= ratio Qg-rQd 339

resistance in ohms 129

revolutions per minute 6

revolutions per second 339

radius of armature 6

resistance per phase of stator - - 418

apparent resistance of secondary referred to the primary circuit - 128

resistance per phase of rotor winding 428

apparent resistance per phase of stator and rotor 428

radius of coil 233

breadth of phase-band 306

number of turns per pole 299

turns of primary 428

turns of secondary 428

width of slot 64

total turns in series 310

turns per coil 306

number of seconds 128

thickness of sheet in cms. 49

thickness of insulated coil 158

thickness of copper strap at right angles to B 144

voltage 141

velocity in cms. per second 306

peripheral speed of armature 479

eddy-current loss in watts per cu. cm. of iron 48

hysteresis loss in watts per cu. cm. of iron 47

no load losses in watts 420

weight of rotor in lbs. 405

l^' 308

2 T

distance from hottest part in cms. 227

apparent reactance per phase 428

Y = apparent impedance 428

y = throw on commutator 612

1 ~-

r2,i =

J3 =
S

^, =

s =

T =

Tc =

t =
/

te =

V =

fa =

We =

Wh =

Wn =

Wr =

X = "

SYMBOI^ 3dx

SYMBOL PAGK

Z = total slots in periphery 309

Za = effective number of conductors on armature ------ 25

Zg = total number of conductors in series 6

Zt = total number of conductors. ZT'r-c=Za 24

a = angular displacement of centre line of pole from uniformly rotating vector 339

a = angle of slope of tip 79

)8 = lu-^Ii 341

y = slot pitch 309

c = base of Napierian logarithms 128

r) — hysteretic constant 47

r\ = amplitude of tooth ripple 316

d = angle of displacement 306

Arf = permeance of body of slot per cm. length of iron 422

X,„ = permeance of mouth of slot per cm. length of iron 422

A^ = permeance of zigzag path per cm. length of iron 424

Ajk = heat conductivity in centimetre measure 221

kh — heat conductivity in inch measure 221

/x = permeability 48

a- — coefficient used in connection with air-gap coefficient . - . . 64

o- = copper space factor 239

o- = displacement on clock diagram showing electrical relations - • - 339

a- = angle subtended by half coil breadth = - - 306

T 2

'^mr^ = flywheel effect 339

T = In^hc 421

T = pole pitch 305

if> = flux interlinking a coil 306

<f> = angle of lag of current 643

<l>g = end-leakage per pole per ampere in stator 426

<f>i = flux leakage per pole across iron teeth per ampere in stator - - - 425

<f>g — leakage flux per pole per ampere in the stator 421

^p — normal flux per pole 421

^ = angle between centre line of pole and current vector . . - . 294

PART I
SHORT RULES

TO BE USED IN THE

DESIGN OF DYNAMO-ELECTRIC MACHINERY

W.M.

CHAPTER I.

INTRODUCTION.

General scope of the book. The term " dynamo-electric machinery " will be
here taken to include: alternating-current generators and motors, continuous-
current generators and motors, and machines for converting from one kind of
current to the other.

It will be assumed that the reader is familiar with the laws of electricity
and magnetism as applied to the design of dynamo-electric machines and that
he is conversant- with the theory and operation of these machines as given in
the many excellent text-books on these subjects.

It has been thought that, amongst the many books on design, there is still
room for one which views the subject more particularly from the manufacturer's
point of view. The problem constantly before the manufacturer is how to
build economically a machine which will fulfil prescribed guarantees. This
book, then, will aim mainly at gi^ang concise methods of designing machines
to meet given specifications.

The ftuiction of the perfomumce speciflcation. In order to treat satisfactorily
of the methods of meeting guarantees, it will be well to deal with the specification
itself and of the conditions of operation which must be kept in vi^w when the
specification is drawn up. There are many different circumstances under which
machines are to be operated. For instance, some alternators are intended to
form part of a small isolated plant and to supply a power load, others to take
the mixed lighting and power load of a large central station: some motors are
intended to work cranes out of doors, others to drive machinery in hot mines.
It is for the user or his consulting engineer in the first place to decide what
characteristics a machine shall have when it is intended to operate under certain
conditions. The question then arises, how should the performance specification
be worded, in order to specify a machine fitted for a particular class of work?
This is a question for the purchaser's adviser. Sometimes the manufacturer
acting in the capacity of advising engineer decides this question.

Secondly, if we have before us a specification and know what the machine
is intended to do, what is the most economical way of building a machine to
comply with the specification and give satisfaction to the purchaser? That is
solely a question for the manufacturer.

4 DYNAMO-ELECTRIC MACHINERY

It is our purpose to consider the various conditions under which each
class of machine may have to operate and to give some typical performance
specifications, drawn up to meet common conditions. The design of a machine
will then be completely worked out to meet each specification, and notes will
be given as to how possible variations in the specification could be met. In
all this we must have regard to commercial requirements and the adherence to
standard rules and to the utilization of standard frames.

Bnles for calctilation applicable to all dynamo-electric machines. Before
examining each class of machine in detail it will be well to deal with certain
matters which are common to all generators, motors and converters, matters
relating to the magnetic circuit, the electric circuit, the insulation, the ventilation
and the framework. A great number of rules, formulae and details of shop
practice on these matters are common to all machines, and it will save time to
dispose of them in a few preliminary chapters.

It is well to have one general method of designing all the machines,
A.C. generators, c.c. generators, induction motors and rotary converters, so
that the experience gained with one class may be readily available for the improve-
ment of another. That such a general method of design is possible can be
seen from the following considerations.

One general method for all machines. All dynamo-electric machines depend
for their operation upon the same fundamental facts — firstly, the fact that when
a conductor is moved across a magnetic field there is generated in it an electro-
motive force, and secondly the fact that when an electric current flows along a
conductor in a magnetic field, the conductor is subjected to a mechanical force.
The calculation, therefore, of any such machine raises such questions as the
following : How much magnetic field ? How much motion 1 How much voltage ?
How much current? How much force? And in addition, we have the im-
portant questions of how much heat is produced and how is that heat carried
away.

Now there are two ways of looking at the fundamental phenomenon of the
generation of electromotiTe force, and these have given rise to two general
methods of design, both of which are commonly used.

According to one way, a certain total flux interlinking with an electric
circuit changes in amount or completely reverses in a certain period of time,
thus generating a certain mean electromotive force in the circuit during that
time.

According to the other way of looking at the matter, a conductor of a
certain length moves in a field of a certain fiuxrdefnsiiy at a certain velocity and
generates for the instant a definite electromotive force.

The first of these aspects of the phenomenon leads us to speak of the total
flux per pole, which we may represent by the letter iV, and our formula for the
electromotive force generated in the windings of the armature of an ordinary
continuous-current generator, in which the number of poles is equal to the
number of paths in parallel through the armature, is

E^nZNx\0-\ (1)

INTRODUCTION 6

where n is the number of revolutions of the armature per second and Z is
the total number of conductors.*

The second of these aspects leads us to speak of the flvx-density in the
air-gap, which we may designate by B, and then the instantaneous value of
the electromotive force generated in one conductor moving at right angles to
the magnetic flux with a velocity of v cms. per sec. is

e = vBlxlO-^, (2)

where / is the active length of the conductor in centimetres.

The first method of calculating the electromotive force has the advantage that
it only deals with the total flux without troubling about the distribution of the
lines of force in the air-gap, but this very feature limits its application to those
cases where we are content to know the mean electromotive force generated in
one alternation of the flux. The formula is therefore not so generally appli-
cable as the second one, which gives us a more complete mental picture of what
is happening under each pole.

Out fdndamental formula for voltage generated. It is possible to have a
combination of these methods which preserves the advantages of both. We may
lead up to it in the following way:

Suppose that we have a rotor surrounded by a stator (see Fig. 1), as in an
induction motor, but the flux in the gap, instead of changing from point to point

Fio. l.-^HomopoIar generator with one conductor.

along the periphery, is all of one sign and distributed uniformly (the return path
being, if we like, along the shaft). Consider the electromotive force generated
in a conductor on the surface of the rotor when it is moved across the uniform
field of the stator. If B is the flux-density in the air-gap in lines per sq. cm.,
r the radius of the rotor in cms., I the length of the rotor iron in cms. and n
the speed in revs, per second, then the total flux passing into the rotor will be
B X 2vrl and the total flux cut per second will be B2Trrln, Writing Ag for the

*In the two-pole case the total ohauge of flux through one turn in half a revolution
18 2N, because the flux changes from + ^ to - ^. In one whole revolution it is 4N, thus
the mean rate of change is 4nN, Now, if Z is the total numl)er of conductors in an ordinary

two-pole drum-wound armature, the number of turns in seiHes is -. Thus we get

E=nZNxl(y\

6 DYNAMO-ELECTRIC MACHINERY

total cross-section of the gap=2irr/, we have the electromotive force E in volts
generated in one conductor,

E^BAgUxlO'^ (3)

or i?=B^^i2p,„x^VxlO-8, (4)

when the speed Epm of the machine is given in revs, per minute.*

Observe that the formula preserves the symbol for the flux-density in the gap,

and that at the same time we have SAg the total flux of the whole frame clearly

before us. The speed is, moreover, given in revolutions per minute instead of

in linear velocity as in formula (2).

The uniform flux distribution, assumed in formula (4), does not ordinarily

occur (except in homopolar machines), but it is possible to apply an equation of

Fio. 2. — Heteiopolar generator with four conductors In aerioB.

this form to any dynamo-electric machine by introducing a coeflicient so chosen as
to allow for the want of uniformity in the flux distribution, and also for any
peculiarities in the arrangement of the conductors. For instance, to take a simple
case, assume that the stator has four poles, each of which has an effective pole
arc only 0'7 of the pole pitch, as represented in Fig. 2. The average electro-
motive force in one cx)nductor would be

E^O'7BAgRp^x^\xlO-K

Now if the flux is not all of the same s^, but changes from positive to
negative as we go from one pole to another, and if there are Zg conductors on
the rotor, joined in series as shown in Fig. 2, the average value of the electro-
motive force will be i? = 0-7B^^,^x^xlO-8xir, (5)

Note that the coefficient 0'7 would be used whatever the number of poles
might be, provided that the ratio of pole arc to pole pitch were the same.

* If we prefer to work in kapp lines per square inch, denoted by Bjr, the formula takes
the ample rorm ^^ BmA;R^ x 10-«,

for one kapp line =6000 c.G.8. lines. Here ^, is in square inches.
Or if we prefer to work in CG.s. lines per sq. inch,

where ^"sarea of the gap in sqwue inches, and B* the flux-density in lines per sq. inch

INTRODUCTION 7

If now the armature current be denoted by I a, the output in watts,

IaE ^0-7 x^ ^10-^ xRj„n^BAgxZjA (6)

The electromotive force coefficient Ke. Now consider that the flux is not
uniform under the poles but varies from point to point, having any value from
0 to B, where B is the maximum value, and that the conductors which are
connected together are out of phase with one another as depicted in Fig. 3.
The same form of equation is still applicable for calculating the electromotive
force, provided we choose such a coefficient as will allow for the peculiarity in
the flux distribution and in the arrangement of the conductors, and also, in an
alternating-current machine, for the taking of the square root of mean square
of the voltage, when the result is to be given in virtual volts. The exact

Fig. 8. — ^Heteropolar generator with varying flnx-denaity and oondnctors out of phaM with

one another.

method of allowing for these things will be given in its proper place. We
wish to point out here a formula of the general form

E:=^KeBAgRjrmZ,X^\xlO-9 (7)

can be used for calculating the electromotive force of any dynamo-electric machine
and that this formula has the following advantages in its favour:

(1) The formula contains the term B representing the maximum value of
the flux-density in the air-gap, and this term, as we shall see later, is useful
in many ways.

(2) The expression BAgy the maximum flux-density multiplied by the total
area of the active surface of the armature, has a fairly definite maximum value
for any given frame, so that if we are familiar with our frame we know by a
glance at the formula to what extent we are making good use of it. For
instance, if we have an armature for an A.C. generator whose diameter is
50 inches and length 10 inches, then -^^ = ir50x 10=1570, and if we know
from experience that B'' cannot be made higher than 60,000 then the maximum
value of B"Al for that frame is 94 x 10«.

As this quantity BAg is almost independent of the number of poles, the
designer soon comes to know the value it should have for any particular
frame, and is able to judge at a glance how far he is utilizing the magnetic
circuit of that frame.

8 DYNAMO-ELECTRIC MACHINERY

(3) The coefficient Ke also has a certain recognized maximum value for a
certain kind of machine. Thus, for a three-phase generator K^ may be equal
to 0'4. If it has a lower value in any calculation under consideration (a&
may be the case where the pole arc is a small fraction of the pole pitch), the
designer's attention is called to that circumstance.

(4) Just as the expression BAg gives us at a glance the magnetic loading
of the frame, so the expression laZa tells us at once the current loading. Here
we have taken /« as the current per conductor and Za for the total conductors
on the armature. If there are a number of paths in parallel, then if I a is the
current at the terminals and Zg the number of conductors in series, the current
loading will be /^^«. In using the method of design given here; the expressions
BAg and laZa are continually in evidence, and we can watch how one decreases
and the other increases in the fight for room which occurs between the iron
and the copper. The output of the frame is of course proportional to the product
of BAg and IaZa»

All the machines dealt with in this book may be regarded as variations of
one type of machine, say, of the alternating-current generator. It may be
said that the differences in the design of the different types of machine consist
in the amount of importance which we attach to certain features. Thus, an
induction motor is a machine with a very great armature reaction, and a very
small air-gap, magnetized entirely by wattless current from the line and provided
with a large amortisseur.

In a generator we attach importance to having a large magnetomotive force
on the magnetic circuit ; while, in an induction motor, we attach importance to
keeping the magnetomotive force as small as possible. In a commutating machine
special attention is given to keeping the self-induction per coil as low as possible,
and preserving a good field form, otherwise inside the armature it is very like
an alternating-current generator.

In fundamental design all these machines are the same, and the formula

E = KeBAgZgR^n X irV x 10"®
is applicable to all.

Methods of calcnlation common to all machines. The calculations of the
magnetic circuits of all these machines are very similar, involving, as they do,
mainly considerations of the air-gap, teeth and iron body of the machine. Again,
the considerations which enter into the calculations of the electric paths are very
similar. The convenient kinds of windings, the calculation of the conductors
and the arrangements for cooling are nearly the same for all the machines. It
is therefore well to take up these general matters in a few preliminary chapters,
and then when we come to consider each class of machine by itself we will be
able to avoid repetition and devote ourselves to those points which relate
particularly to that class.

Judicious gnessing. It must not be supposed that the rules given in the
subsequent chapters are intended to be employed in all cases in which they are
applicable. A busy designer would never get through his work if he stopped to
calculate everything. He guesses a great deal, or makes rapid mental estimates
of quantities he has not time to calculate. Now, he is never justified in so

INTRODUCTION 9

guessing unless he knows the limit of his possible error with fair accuracy, and
knows that with the error he will still have a machine which will comply with its
specification* Knowledge of these two things can only, come from many calcula-
tions made and many machines tested. The way to acquire the art of correct
guessing is to employ fairly simple rules for calculation that are based on sound
principles. An empirical rule, however often applied, does not help the mind to
form rapid mental estimates, because it does not take into account all the factors
that determine the result.

While some of the rules here given may seem to lead to calculations which
are too lengthy for ordinary shop use, it must be remembered that an hour's
calculation may sometimes save the designer weeks of worrying experience. The
great art is to know what to calculate and what to guess.

CHAPTER II.

THE MAGNETIC CIRCUIT.

Field-form and field-form coefficients. We shall assume that the reader is
acquainted with the laws of magnetism and their application to the design of
dynamo-electric machines. Our object in the following chapters will be to collect
for his convenience rules which are useful in the calculation of magnetic quantities
in the commercial design of machines and to emphasize those points in the magnetic
design which experience has shown to be of importance. At first we will only
consider those points which are common to all electrical machines whether for
alternating or continuous-current.

THE EFFECT OF THE NUMBEE OF POLES ON THE GENERAL DESIGN.

Different numbers of poles on an armature of giv«n diameter. The fixing of
the number of poles which a machine shall have is one of the matters to be taken
up later, when we are considering each machine in its own class ; but we may here
look at the general effect on the design of having few or many poles, irrespective
of the question whether the machine is for alternating or continuous-current or
whether it is a generator or a motor.

In the first place, we know that for a given speed, given ampere-wires per inch,
and given fiux-density in the gap, the output of a machine is proportional to Z>^Z,
where D is the diameter of the active face of the armature and I the length of the
iron. As a first approximation, it is independent of the number of poles. Now,
if we fix D we can draw a circle which represents the periphery of the active face
of the armature, and we can draw out diagrammatically the magnetic circuits for
a two-pole, a four-pole, a six-pole and an eight-pole machine, as is done in Figs. 4,
5, 6 and 7.

In these figures the outputs are supposed to be the same at the same speed.
The diameter is constant, and for the moment we will take the air-gap the same
in all cases (though in practice it would usually be greater when the poles are
fewer). It will be at once seen from these diagrams that, under the conditions
specified, the machine with the few poles requires more iron than the machine
with many poles. The dimensions a and h are supposed to represent the depths
occupied by the teeth and the windings, and the dimensions c and e are the depths

THE MAGNETIC CIRCUIT 11

of the iron behind the bIoIb which serve as paths for the magnetic flux. In the
two-pole case (Fig. 4) the magnetie flux which threads through the rotating
element has only two paths hy which to return, and the depth c must therefore
be made very great. Moreover, if the density in the gap ia reasonably great, wa
will require the whole of the radius e to carry the flux. Where the rotating

Fia. s.

orUag dlumtoc

element is a field magnet we can utilize the steel of the shaft to carry the flux,
but where the flux is alternating the dimension e must not include the shaft (see
Fig. 8), BO that in the two-pole case we would be under an additional disadvantage,
for we have to increase D in order to make room for e. This still further increases
the total quantity of material in the machine.

IHagmiiniKUc vlcwt of ■ix-polg utd dihi-pole icmnton ol the unw working diameter,
Bhoouic the nlaUve MnouDte ot Iroa rcqnlnd.

In the four-pole case (Fig. 5), assuming the same flux-density in the gap, the
depths c and e need only be about one half as great as in Fig. 4. It should be
remembered, however, that where the frequency is doubled (say 50 cycles instead
of 25) it is usual to work the iron at a rather lower density.

In the six-pole case (Fig. 6) the iron behind the slots is still further reduced,
and in the eight-pole case the machine assumes the general proportions indicated
in Fig. 7.

12 DYNAMO-ELECTRIC MACHINERY

It ia not only in the magnetic circuit that the two-pole machine takes more
material than the four-pole, and the four-pole more than the aix-pole. In the
electric circuit also the end connectiona are longer and more bulky when the
poles are fewer. In the atmve figures we have taken the speed constant, and
ths frequency therefore increases with the number of poles. The result is as we
would expect ; there is less material required at higher frequencies.

In those cases where the freqiiencij is p-esciibed and the speed may be chosen,
it usually pays to adopt a six-pole construction in preference to a four-pole, not-
withstanding the fact that the speed is lower in the six-pole case. The material
required for a four-pole 'J5 cycle machine running at 750 R.P.M. is rather less than

uvtng the ume ootpot, at U>e ume

for a two-pole machine of the same output running at double the speed. There
may, however, be good reasons for adopting the higher speed, as, for instance,
where a steam turbine is used for driving and the higher speed gives better
economy.

With continuous-current machines, where the speed is presciibed and the
frequency may be choseu, the number of poles sometimes depends on the desirable
number of brush arms, but apart from this consideration one will not adopt a
two-pole construction in preference to a four-pole construction unless the size is
so small that the reduction in the cost of labour is more important than the
reduction in the cost of material. The difference in the amount of material for
the two-pole and the four-pole cases when the machine has inwardly projecting
poles can be seen at once from a glance at Figs. 8 and 9. It is even more striking

THE MAGNETIC CIRCUIT 13

than the cases considered in Figs. 4 and 5. In these cases, as the machines are
supposed to be of the same output, the same speed and the same length of iron,
it has been necessarj to increase the diameter of the two-pole machine in order to
make room for the shaft, which cannot carry alternating flux. The provision of

sufficient cooling surface on the two-pole field coils necessitates either a very long
pole limb or a great depth of winding on each pole. But these are matters which
will be more properly considered under their proper headings.

THE FIELD-FORM.

Dixtribation of magnetic flnx in the air-gap. The value of the co-efficient K,
in the equation,

depends {mltr alia) upon the way in which the magnetic flux is distributed in the
^p. We will consider at this point how the field-form may be conveniently
plotted and how the coefficient K, may be determined for various types of

There are two classes of cases to consider, (1) where the magnetomotive force
is created by a coil on a simple salient pole as in ordinary continuous-current
machineB and engine-type alternators, and (2) when the magnetomotive force is
supplied by a number of coils distributed over the pole face.

Fisld^fonn nnder & salient pole. In the case of the simple salient pole, we
have a definite difi'erence of magnetic potential between the pole and the armature
iron, so that the density in the gap at any point depends mainly on the length of
the gap at that point, and where we can neglect the saturation of the iron parte, it
1b inversely proportional to the length of the gap.

As an example, let us plot the field-form of a rotary converter in which the
Armature teeth are not highly saturated. Let the pitch of the poles be II"

DYNAMO-ELECTRIC MACHINERY

measured on the periphery of the armature, the width of the poles 8", the air-
gap being 0*2". Let the pole have a 1" bevel such that the air-gap at the
corner is 0'3".

Take a sheet of squared paper. Near the bottom draw a horizontal line to
represent part of the periphery of the armature (Fig. 10). Choosing some con-
venient scale, draw two vertical lines, one for the centre line of the pole and one
for the neutral line, as shown. Then draw in half of the pole and air-gap, as
shown by the line EST, Choose some convenient scale of ordinates for the flux-
density, taking the maximum density in the air-gap as unity (the actual values of

no. 10. — Field-fonn diagram for a saUent pole with no saturation.

the density do not concern us for the moment). To construct the field-form, draw
through ordinate 1 a horizontal line AB, the same length as ES, extending in this
case 3 inches from the pole centre-line. Here the pole bevel begins. If we
neglect for the moment the fringing effect, the flux-density on the surface of the
armature, under the comer of the pole T, would be 0*66. Similarly, half-way
along the bevel it is 0*8 Plot these two points, and draw the curve through them
as shown. Now we must consider the fringing. The shape of the fringing curve
depends mainly on two things, the length of the air-gap g at the corner and the
distance c from the comer to the neutral line. Mr. F. W. Carter has given "^ us a
method of calculating the fringing curve for any given value of c/g, and has pointed
out that where we are given the fringing curves for several values of c/g we can
draw by eye the curves for intermediate value with sufficient accuracy for practical

*Jour, Ifut. Elec, Engrs,, 1900, part 146, vol. xxix. ; Eltc, World and Eng., Nov. 90, 1901.

THE MAGNETIC CIRCUIT

purposes. In Fig. 1 1 are reproduced the curves which Mr. Carter has plotted for
c/g = 2-5, cjg^b and c/<7 = infinity. In the case of a pole with a slight bevel, such
as shown in Fig. 10, the distribution of the flux in the interpolar space will be
almost the same as if the air-gap were uniform and of the same length as at the
corner T, in this case 0*3 inch. For the purpose of drawing the fringing curve
we must take the ordinate 0*66 as if it were the full value 1 in Fig. 11, and
all other ordinates must be taken in proportion. Taking cjg = 5, in our case, we
could, if we liked, plot the curve shown in the dotted curve and so complete the
field-form ; but our object is not so much to plot the exact shape of the field-form
as to find its area, and it is possible to find the value of the flux between the

Fio. 11. — Corves showing distribution of fringing flux for different values of ejg.

comer and the neutral line in a more convenient way. Imagine that the pole is
made wider at each side by a certain amount, Kag, such that the flux in the gap
under the added part is just equal to the fringing flux. Mr. Carter has given us
the value of the coefficient K^ for diffierent values of cjg. These values are given
in Fig. 12. Two curves are given,; ^ and B. The curve A relates to the case
where the pole has a square corner and the flank of the pole is approximately at
right angles to the surface of the armature. The curve B relates to the case
where the pole is provided with a spur of the shape shown in the sketch at the
side of the figure, there being an angle of 135 degrees between the side of the
pole spur and the surface of the armature.'^ For any intermediate case it is
easy to judge with sufficient accuracy the position of an imaginary curve drawn
between A and B, Instead of plotting out the fringing curve, all that is necessary
is to set off DQ, as shown in Fig. 10, and complete the flux curve as shown by the

* Messrs. Hawkins and Wallis in their excellent book on the dynamo (page 449 of the 1909
edition) give curves for various values of the angle between the pole and the surface of the
armature.

DYNAMO-ELECTRIC MACHINERY

full line. We obtain the length of DQ as follows: Take the ratio of c to g^ — in this
case 5. From Fig. 12 curve A, abscissa 5, gives us Ka=^ 1*25. Make DQ= r25<7,
in this case 0'375". The area under DQ = area under curve DN, The area of the
figure OABDQM is proportional to the flux from the half pole, and the ratio of
this area to the area of the figure OAPN gives us a certain coefficient, which we
will write Kf, the/iw; coefficient If the maximum flux-density in the gap extended
for the whole pole pitch, the flux from the pole would have a hypothetical maximum

f — . * *#■

1
1

(

'^ase A

— rxr-

t-9-

/a-

V///A.

'"/M.

/•7-

/■e-

— 15-
13-

y//A

y/y/y.

V///'

y////.

"^ '

^ —

1
1

a7-
— M-

04-

Case
Y///A

^
V •

\ 1

V 1

f 1

O-f-

y//A

0 / ;

fames of

1 -^

( i

7 6 0/

0 i

t It i

i k a

Fig. 12. — Values of frlDglng coefficient Ka tor different Tallies of e/g both for case A

and case B,

value corresponding to twice the area of the rectangle OAPN. The coefficient K/
is the coefficient by which we must multiply this hypothetical maximum flux in
order to get the true value of the pole flux.

Working upon squared paper, the figure can be sketched with great rapidity by
hand, and taking the value of Ka from Fig. 12 we easily and accurately make
allowance for the fringing flux. To get K/^ the easiest way is to run the stylus of
a planimeter'^ around OABDQM, and then again around OAPN. The ratio of
the readings gives us K/. In Fig. 10 ^=0*738.

* Every dynamo designer should have a planimeter at hand, because by means of it he can
make quick and accurate estimates of quantities he otherwise would not take the trouble
to calculate. A good plan is to work on a drawing board upon which a sheet of tracing cloth
is always stretched. If the area of any figure is required, a sketch of the figure is made with
a soft lead pencil, the sketch is put under the tracing cloth which serves to hold it in position
and the stylus of the planimeter is run along the penmeter without any fear of that vexatious
catching of the edges of the paper asainst the planimeter which sometimes happens when the
size of the paper is not much bigger tlian the figure.

THE MAGNETIC CIRCUIT

Cases where the saturation of the teeth caxmot be ne^rlected. Where the
saturation of the teeth is fairly high, as in the case of continuous-current gene-
rators, allowance must be made for it, or the value of Kf obtained will not be
sufficiently accurate.

The method of allowing for the saturation is really a method of trial and error,
but where we know beforehand, as we generally do, the fraction of the total
.ampere-turns per pole which are to be expended on the teeth, we can get at Kf
with fair accuracy by the following construction :

Consider a continuous-current generator with a pronounced pole spur (Fig. 13),
the side of the spur VT making an angle of about 135 degrees with the surface of

Fio. 18. — ^Field-form diagram for a salient pole with oonslderable saturation Id the teeth.

the armature, and the length of VT being about the same as the length TN, This
is an average case among modem generators. Even where the dimensions differ
considerably from those given, the method here described will give fairly accurate
results.

First neglect the saturation of the teeth, and draw the figure OABDQM as
before, except that instead of obtaining a value Ka from the curve A in Fig. 12,
obtain it from the curve B. In the case illustrated, c/g=^b and Ka therefore = 1.
We make i)Q=l xg and complete the figure OABQM.

Now, it will be seen that the effect of the saturation of the teeth will be to
diminish the flux-density under the pole, while it does not affect to any great
extent the distribution of the flux between the poles.

»» • M»

18 DYNAMO-ELECTRIC MACHINERY

The presence of the teeth will cause the field-form to have ripples on it which
move forward with the armature. In practice it is not worth while to take
account of these ripples. The field-form here plotted may be taken as the
average field-form with the ripples smoothed out.

Suppose that we intend to expend 20% of the total ampere-turns on the
pole in driving flux through the teeth, which are in this case somewhat saturated.
Instead of having, say, 5000 ampere-turns expended on the gap, we will have
available only 4000, and this will have an effect upon the general shape of the
field-form, because the fringing will be about the same as if there were 5000
ampere-turns, while the flux under the pole will be diminished to 80 % of what it
otherwise would be. We accordingly proceed as follows : Through the ordinate
08 we draw the horizontal line A'B". This gives us the top of the corrected flux-
distribution curve. Having regard to the amount of the bevel and the reduction
in the saturation of the teeth which will occur under the corner of the pole,
roughly estimate the fraction of the ampere-turns expended on the teeth under
the comer of the pole. This in our case may be about 0 05. Take D* accordingly
0*05 (on the ordinate scale) below D and complete the curve B'jy by hand. Join
lyQ. Observe that the allowance for the fringing flux (the part of the figure
under nQ) is almost the same as if there were no saturation in the teeth. Con-
tinue A'B' to jP. Then the value of K/ is obtained by finding the ratio of the
area of the figure OA'B'D'QM to the area of OA'FN, Observe that the value of
the Kf is greater in this case than if we had taken the ratio of the area OABDQM
to the area OAPN, so that the saturation of the teeth increases K/ for a given
shape of pole.* The value of K/ in the case given in Fig 13 is 0*75.

One of the advantages of this method of working is that it enables the designer
without any elaborate calculations to make allowances for minor matters aflecting^
the distribution of the flux. Suppose, for instance, that the pole tip is highly
saturated. If we know approximately the number of ampere-turns expended on
the pole tip, we can allow for it in Fig. 1 3 by putting ly lower down, just as A' ia
put lower down, to allow for the drop in the armature teeth.

Field-form under a distributed winding. When the diflerence of magnetic
potential between armature and field-magnet is not uniform all along the surface
of the pole, as where the ampere-turns are supplied by a distributed winding, the
first step is to make a diagram to give us the distribution of magnetomotive force.

Take the case of a four-pole cylindrical field-magnet whose coils lie in slots on
the pole face, such as is illustrated in Fig. 371, p. 400. Let there be 96 slots,
80 wound, 4 slots being left vacant at the centre of each pole. Take the diameter
at 36" and the diameter of the bore of the armature at 37|'', thus the radial gap
is f'. It is best to measure the pole pitch, not on the surface of the cylinder,
but half-way across the gap, that is on a circle whose diameter is 36| inches.
The pole pitch is 28".

Theife are 24 slots per pole. Lay out on squared paper a horizontal line having
25 divisions, numbered 0 to 24 (see Fig. 14) ; these represent 25 teeth, and the

* It must be remembered that the height of the ordinate OA is immaterial. All that we
want for the moment is the ratio of the two areas named. The voltaee of the machine will
then be a function of the maximum density in the air-gap and the coefficient A/.

THE MAGNETIC CIRCUIT

spaces between these give us 24 slots. Mark oif the 4 unwound slots in the centre
of the pole. Mark off (or imagine marked oif) the end connections, connecting
slots 10 and 15, 9 and 16, etc. There are five teeth in the centre of the pole upon
which the full ampere-turns of the coils are exerted. Represent the full ampere-
turns by the ordinate 1, drawing the line AB over the five central teeth. The
ampere-turns on the remainder of the teeth are less and less as we get further from
the centre, and can be represented by the sloping dotted line in Fig. 14. The
magnetomotive force really increases in steps but it is not worth while to take
account of these. Now, if there were no saturation in the teeth the field-form

FiQ. 14. — ^Field-fonn with a dietribated winding and saturated teeth.

would have the same shape as the magnetomotive force curve ; but in cylindrical
field-magnets it is usual to saturate the teeth until they require for their magnetiza-
tion a considerable percentage of the total ampere-turns. Suppose that it has been
decided to expend 20 % of the ampere-turns on the central teeth. We can draw
a horizontal line A'B, having an ordinate 0*8 to represent the flux from the five
centre teeth and complete the figure OA'FN, which is of ^ such a form that any
portion of an ordinate, such as A A' or CC, represents the fraction of the magneto-
motive force taken to magnetize the teeth.

The exact shape of the curve OCA' will be considered when we come to deal
with the ampere-turns on the teeth. "^ It can, with a little experience, be drawn

* See page 78.

20 DYNAMO-ELECTRIC MACHINERY

by hand with sufficient accuracy for the purpose of getting £/. The ratio of
the area OA'BN to the area OTFN gives us K/. In the case taken in Fig. 14,

THE FIELD-FORM OF INDUCTION MOTORS.

Strictly speaking, the field-form problem in an induction motor is inverted.
Instead of starting with a certain magnetizing current, and then building up the
field-form, and from that the electromotive force wave-form, we should, to be
logical, start from the wave-form which is impressed upon the motor and work
backwards to the magnetizing current. If we could do so, we should find that the
current would in most cases not be at all sinusoidal, and that not only would the
third harmonic be very pronounced on account of a saturation of the iron, but
there would be numerous harmonics introduced to suit the particular kind of
winding in the stator slots and generate in it the electromotive force like that
impressed upon the machine.

Even if we knew the wave-form of the electromotive force that will be
impressed upon the motor, the problem of finding the exact form of the mag-
netizing current would be extremely difficult. The usual practice is therefore to
assume a sine wave-form for the magnetizing current and give it sufficient ampli-
tude to create the flux required to generate the electromotive force of the motor.
The designer knows that he is here making an unwarrantable assumption, and he
is not surprised when the wave-form of the real magnetizing current rather upsets
the calculation of the power factor of the motor. Going then into the problem
with our eyes fully open to the defects in our method, we can proceed.

Take a three-phase motor, with three slots per phase per pole and a full pitch
winding. Lay out the air-gap in a straight line and mark off the stator slots as in
Fig. 15. It is not of course necessary to draw the slots. Assume first that the
magnetizing current is at its maximum in phase A, and at half its maximum in B
and C.

The numbers 1, 1, 1, 0*5, 0-6, 0*5, etc., along the top of Fig. 15 are propor-
tional to the ampere-wires in the slots immediately above them. The tooth
between the slots 6 and 7 has the maximum ampere-turns upon it, and on the
middle of slot 2 lies the centre point of the band of magnetizing current.

In order to get an idea of the field-form of an induction motor, first lay out
the magnetomotive force curve ODEFGHIJNy beginning under the centre of
slot 2 and having its maximum under the tooth 6, 7. Under the centre of slot 2
the vertical line of height 1 is bisected at 0, because the ampere-turns in slot 2
may be said to be half on the pole to the right and half on the pole on the left
Under slot 3 the magnetomotive force curve rises by an amount 1 and under slot 4
by an amount 0*5 and so on, until we get down to the point iV. Before we can
draw the field-form we must know what percentage of the ampere-turns on the
pole are expended in magnetizing the iron ; this percentage could only be arrived at
properly by a method of trial and error, but for practical purposes it is sufficient
to take a percentage which the designer finds most economical. In Fig. 15 we
have taken 23 % of the ampere-turns per pole as expended on the iron. Thus the
maximum ordinate of the field curve is only 77 % of the magnetomotive force

THE MAGNETIC CIRCUIT

curve. At lower flux densities the saturation is not so great and the field-form
follows more nearly the curve of ampere-turns, as shown by the thick line.

To plot this field-form more accurately it is necessary to work with an air-gap-
and-tooth saturation curve as shown on page 78.

If we take the distribution of the magnetizing current after it has advanced 30**
in phase, the ampere-wires in the various slots will be 0*86, 0*86, 0*86, etc., respec-
tively, as given in the second line along the top of Fig. 15, and magnetomotive

1 — 1

A
2

B
4

C
7

J. ..t

c
s

0
0

A
U

a
-/

•46

'S6

1
1

- "1/

-.«

f'^' M

■ ^ ■/

•

■

- *

^ f—-

\L_

-1"
■V j

1 ^^^

l:^

"/I

•

1
L_

— I

./^

^ 1
1

•

■

Fio. 15. — ^Field-fonn of induction motor for two different shapes of magnetomotlTe-foroe

wave-form.

force curve will be of the form shown by the chain dotted cun^e KLMFQBS.
The field-form will then be slightly different in shape, as shown by the thin full
line.

If we run a planimeter around the flux curve shown by the thick line, and then
around the rectangle whose height is equal to the maximum height of this curve
and whose base is given by the pole pitch, we will find that the ratio between the
readings is 0*68. This is therefore the value of K/ at the instant when the
magnetomotive force is as shown by the dotted line ODEFGHIJN, If we go
through the same process for the thin-line curve, we will find that the value of Kf
comes out 0*695 for the instant when the magnetomotive force has a distribution
as shown by the line KLMFQBS. The area of the thin curve is 5 % less than

22 DYNAMO-ELECTRIC MACHINERY

the area of the thick curve, so that the highest ordinate in the thick curve is
proportional to the maximum B^ and the constant 0*68 used in connection with
this flux-density will give us the maximum flux per pole.

The field-form of any machine can be worked out in the same way as shown
in these examples.

In this chapter and the next we give some simple graphical methods of laying out the
field-form and the E.M.F. wave-form, and for calcalating the value of K^ It may be well
here, and in some subsequent notes, to give the analytical methods by which the wave-form
of the B.M.F. can be calculated. As a first step, it is necessary to express the distribution of
B in the air-gap by means of Fourier's series. For a symmetrical curve on no-load this

^H ^' B,= Bi8in^,+ Bjsin3^,+B8sin5d,-i-etc., (1)

where B^ is the flux-density at a point x on the periphery of the armature, and Bx is the
angle on a two-pole machine which x has passed through, measured from the neutral plane
(see Fig. 321, p. 305).

Where the field-form is a simple rectangle^ (see Fig. 322), we have

Bx=- Bp ( cosasin^x-f ^ cos3a8in3^ + ? cosSa sin5^x + ... ] (2)

Where the field-form is a trapezium (see Fig. 323), we have*^

and writing /J=^.

B,= ' { sin o sin ^,+ Q sin 3o sin Z0x + etc. j,

.(3)

B,=^ ^ ^sini9^sin^,+|sin3/5|sin3^,+ ...j (4)

Where the field-form is not of any simple shape, these coefiicients, Bj, Bg, B3, etc., can
be determined by any of the methods of harmonic analysis, f

The wave-form of the k.m.f. generated in a band of conductors, moving with a velocity
V at right angles to the direction of B, will depend upon the width of the phase-band of
conductors and their arrangement in slots. When the field-form does not follow a simple sine
law, the placing of the conductors in slots may give rise to ripples in the wave-form of the
B.M.F., and calls for very careful analytical investigation if a more exact wave-form is to be
ascertained (see pp. 304 and 313). The effect of these ripples in changing the virtual value of
the B.H.F. generated in three-phase generators is, in practice, usually very small. Even
where a ripple is very noticeable on an oscillogram, its effect on the virtual value of the
voltage will be small, because the vector representing its maximum value must be added at
right angles to the fundamental vector (see p. 33). We are therefore justified, in the
graphical method given in the next chapter, in neglecting the effect of the high-frequency
ripples in calculating JT*.

*Dr. S. P. Smith, "The Non-salient Pole Turbo Alternator and its Characteristic,"
Jour, Inst. Elec. Engineers^ vol. 47, p. 562. See also paper by Smith and Boulding, xhvi.
Jan. 1915.

t Fischer-Hinnen, Elektrotech. Ztii,^ xxii. p. 422, 1901; EhhtroL tc MaschinevJbau, xxvii.
p. 335; and see Silvanus P. Thompson, Proc, Phys, Soc., xix. p. 443 (1905), Ehctriciany
Iv. p. 78, and Proc, Phys, Soc., Aug. 1911, p. 334; R. B^ttie, Electncian, Ixvii. pp. 326,
370, 444 (1911). See footnote ibid,, p. 326, for list of references to literature on the subject.

CHAPTER III.

THE MAGNETIC CIRCUIT (continued).

THE ELECTROMOTIVE FORCE COEFFICIENT K,.

We will now proceed to give the methods of determining the constant Ket by
which the electromotive force of any machine is calculated when using the formula,

E^^K^x revs, per sec. x conductors in series x AgB x 10"®.

The calculation of the coefficient Ke. In commutating machines of the ordinary
type, in which the pitch of the armature coils is approximately the same as the
pitch of the poles, the coefficient Ke is the same as the coefficient K/. The
reason is that the electromotive force generated in the conductors of a machine of
this kind is proportional to the average value of the flux-density in the gap. In
s, continuous-current machine having a field-form like that given in Fig. 13
with a coefficient, A/ =0-75, the electromotive force generated in all the conductors
in series between two brushes is 0*75 of what it would be if the flux-density were
uniform all along the gap and of the same value as the maximum flux-density in
Fig. 13.

Therefore, in a continuous-current machine or rotary converter, where we are
given the constant K/ for the field-form, we have at once the constant Kg for
finding the electromotive force.

Example 1. The diameter of the armature of a certain frame is 25'' and its length 11", so
that the area of the active surface Ag = irDl=S^ sq. in. Assume that the flux-density in the
gap is 60,000 lines per sq. in. Then the magnetic loading, A^B" is 5*2 x 10^. How many con-
ductors must we have in series on the armature, to generate 500 volts, when the machine is
running at 900 revs, per min. ?

First find the field*form constant K/, Let this be 0*7, then Ke=0'7. From page 6 we have

600=0-7 X Y^ X 5-2 X 107 X 10-« X iT.,

Z.=92.

If 92 is not a very convenient number of conductors to get into the armature slots, we might
-choose the number 96, and make 12 slots per pole with 8 conductors per slot. We would then
oheck over our calculations again as below,

500=0-7 X ^^- X ^;B" X 10-8 X 96.

ffr^ff

This gives us -^I'B' =4-96 x 10^.

24 DYNAMO-ELECTRIC MACHINERY

£xAMPLE 2. Suppose that we wish to build a rotary converter running at 500 R.P.M. to
generate 560 volts on a 6-pole frame, having an armature diameter of 36". We wish to have
54 commutator bars per pole, giving 108 conductors per pole in a lap winding. What will
be the length of the armature if the flux-density is not to exceed 10 kapp lines per square inch
in the air-gap? Having found (see page 14) the field-form constant £/=0'73, ^Tite

560=0-73 X 500 X 108 X Jl^'B^ x lO-^,

^;Bjf= 14200.

If Bk=^0, A = W20=irxdxL

Now rd = 113; .'. ;=12-5.

Example 3. A small motor is running on a 250 volt circuit. Diameter of armature = 28 cms. ^
length 14 cms., speed 1000 B.P.H. Total conductors, 384 in two-circuit winding, giving 192 in
series. What is the flux-density in the gap ? Allow 10 volts drop in armature and brushes^
giving the back E.M.F. =240. Let K/=0'72.

240 x 10^ = i:, X Rp„ X B X J[^ X Z, X ^ (see page 16).

Now ^^=Tx28x 14=1230;

/. 240 X 108=0-72 X 16-6 x B x 1230 x 1»2,
B=8500.

The calculation of Ke for an alternating-current machine. In alternating-
current machines it is convenient to arrange the coefiScient Kg so that it makes
provision for the fact that the voltage measured at the terminals is the square-
root of the mean square voltage, and also for the particular arrangement of
winding whether two-phase or three-phase.

In actual practice Kg has usually been determined once and for all for the
type of field-magnet employed, and it is very seldom that the designer of a
machine has to go through the process of determining it. Nevertheless it i&
well for him to always bear in mind the factors upon which Ke depends.

Let us take an ordinary three-phase star-wound generator, and calculate Ke
for a given field-form. The voltage measured at the terminals of a three-
phase star-wound generator is the resultant of the electromotive force generated
in all the conductors in two legs of the star. These conductors are distributed
over an arc of 120 degrees, and there is therefore a wide difference of phase
between the electromotive forces generated in the first and the last conductors
of the phase band. Our method of calculating Ke must therefore take into
account these differences of phase as well as any peculiarities of the field-form,
and also the fact that only two-thirds of the whole armature conductors are
in series between the terminals.

It is convenient to take the symbol Za in the formula for a three-please
machine to represent the total number of conductors on the armature (except
of course where there are several paths in parallel, in which case Za would be
equal to Zr-^-Cf where Zt is the total number of conductors and c is the
number of paths in parallel).

We will therefore give Ke such a value that the virtual volts

E^Kexrevs. per sec. xZaxAgBx 10"^ (1)

Or in kapp units,

E = K^xu,v.i^\.xZ^xAlBgxlO-^ (2)

THE MAGNETIC CIRCUIT 2^

Or in C.O.S. lines per sq. inch,

E=^K^y. revs, per sec. y^Z^y. A'JSi' x 10"^.

In all the formulae K^ is the same. Ag is the area of the active face of the
armature in sq. cms. = iri>/. A"^ is the area of the active face in sq. inches.

The value of K^ depends not only upon the field-form, but also upon the
arrangement of the armature conductors. The simplest three-phase case to take
is where the field-form is sinusoidal and the armature conductors are distributed
uniformly, each phase occupying exactly 60"^ of arc, the phases being connected in
star in the usual manner (see Fig. 116, page 97). We then have f of the conductors
in series between any two terminals. The breadth coefficient is the ratio of the

chord of 120* to its arc or 1*73-!-^ = 0-828. We can therefore in this simple

case calculate the value of K^ directly.

We have ^« = | x 0-828 x 0-707 = 039.

Similarly for the simple two-phase case with sinusoidal field-form,

^,= 1x0-9x0-707 = 0-317.

It is well to remember these two numbers, as they give us an easy check
on calculations of Kt when the field is not sinusoidal. For instance, in an
ordinary three-phase case, where the field-form is rather broader than the sine
wave, we expect to get Ke rather more than 0-39, and where the field-form is
more slender K^ will be less than 0-39.

We will work out the constant K^ for a given flux distribution in a simple
star-connected armature. In this case Z^ stands for all the conductors on the
armature, unless there are two or more conductors in parallel per phase.
If there were c paths in parallel per phase, we should have to divide the
total conductors ^r by c to get Zq,, Similarly in three-phase mesh-connected
armatures and in two-phase armatures, unless there are several paths in parallel
per phase, Za represents all the conductors on the armature.

These symbols we will use throughout the book.

W^e will later give some curves from which we can read off the values
of Ke for different shapes of pole, but it is well to see how K^ is calculated
in any particular case.

Take for example a three-phase alternator having a pole with the bevel
shown in Fig. 16, the pitch of the pole is 14 in. the width 8 in. The air-gap
is 0-4 in. and there is a one-inch bevel on the corner of the pole, increasing
the air-gap from 0*4 to 0*6 in.

First plot the field-form in the manner shown in Fig. 10. We will assume
that there are 12 conductors per pole, that is 4 conductors per phase per pole.
These conductors are shown by the little circles. Write down the values of
the ordinates of the flux cun'^e at points over the equally spaced conductors,
as shown in the figure. It is best to place the conductors relatively to
the pole so that these ordinates may fairly represent the average flux
immediately adjacent to the conductor. It will be seen in Fig. 16 that if
we take the ordinates 0, 9, 28, 80, 100, 100, 100, 100, 100, 80, 28 9, 0,.

DYNAMO-ELECTRIC MACHINERY

they represent suflfieiently well the distribution of the flux. Here we have
taken an arbitrary value of 100 for the maximum ordinate of the field.

Now, we know that in a star-connected armature the voltage between the
terminals A and C is generated in the conductors of the phases A and -Cm
series with one another. Consequently, when the conductors are in the position
shown in Fig. 16, the instantaneous value of the E.M.F. generated in A and C will
be proportional to the sum of all the eight ordinates 0, 9, 28, 80, 100, 100, 100, 100.

FUld Form

Fio. 10. — Field-form of three-phfue generator, showing relative values of flax-density

opposite the conductors of the three phases.

When the pole has moved to the right over the pitch of one conductor, the instan-
taneous E.M.F. will be proportional to the sum of 9, 28, 80, 100, 100, 100, 100
and 100, and so on. Consequently we may find values which are proportional to
the instantaneous values* of the e.m.f. at various instants throughout the cycle, by
the process worked out below. The process is as follows : From the sum of eight
ordinates subtract the ordinate on the left and add a new ordinate on the right.

* The method g ven here does not enable one to plot the exact wave-form of the electro-
motive force as it would appear on an oscillograph. Where a slotted armature is employed
there is a continual change in the field-form as the pole moves in the vicinity of the armature
teeth, giving rise to high-frequency electromotive forces, which are superimposed as ripples
upon the main wave-form of the electromotive force (see p. .309). In modem mechanics these
ripples are avoided as far as possible by bevelling off the corners of the pole, by making the
slots semi-closed and by employing a sufficient number (not less than six) of slots per pole. Where
proper precautions of this kind are taken the ripples are of little consequence, and the dis-
turbances in the field- form do not affect the virtual value of the generated electromotive
force.

THE MAGNETIC CIRCUIT

The values which are obtained after each operation are distinguished by being
«nclosed between heavy lines.

100

|517 1

subtract 0

Squared
orainates.

267,000
380,000
474,000
474,000
380,000

|408 1

100

308
-28

|280|

100

180
-80

Squared
orainates.

78,500

517
add 100

|617 1
subtract 9

1 ^^1
100

0
-100

10,000

608
add 80

l-iool

-180
-100

10,000

1 688 1

subtract 28

660 •
add 28

|-280|

-308
-100

78,600

f688|

subtract 80

608
add 9

|617 1
100

i -408 1
9

-417

-100

167,000

517
0

|517 1

100

417
-9

1 -517 1

267,000

267,000
167,000

|408 1

These values repeat themselves through successive cycles, and the best check
on the arithmetical process we have gone through is to see whether we have come
back to the same value for the sum when we have come back to the same relative
position of conductors and pole. Thus, in the example given, we start with 517,
and after twelve operations we get back to 517, but the value is now negative,
because we are under the pole of opposite polarity. If we now plot the values we
obtain a curve like the thick curve given in Fig. 17. This gives us the e.m.f.
wave-form of the alternator. It is best to begin this plot where the values change
from negative to positive, because between the positive and negative value there
will be a point where the B.M.F. is zero. It will be noted that though the field-
form is often angular and very far removed from a sine wave, the B.M.F. wave-form
has its comers more rounded off, because it is really the summation of eight field
forms, each displaced by one-twelfth of the pole pitch (see pp. 33 and 304).

The next step is to find the square root of the mean square value of the E.M.F.
wave-form. For this purpose square each ordinate, and plot again as shown in

DYNAMO-ELECTRIC MACHINERY

the thin curve in Fig. 17. If we were to run a planimeter around the curve thus-
obtained, and divide the area by the length of the base, we would obtain the mean
value of the square, and the square root of this would give us the square root of
mean square. But it is not necessary when finding K^ to trouble about the length
of the base line. We argue in this way. If all the twelve conductors were con-
nected in series, and if the field-form were a rectangle of height 100, extending
along the whole pitch of the pole, then the maximum B.M.F. generated would be
1200 as against 700, as given in Fig. 17. Moreover, the E.M.F. would remain at
1200 all the time. The square of 1200 is 1,440,000. This taken as an ordinate
in Fig. 17 would be off the paper, but we can plot it to one-tenth scale, as shown
by the dotted rectangle.

60MCe

SCO

-//

800

700

600

/ >

SCO

& 400

■z

^300

-A

200

r— —

1
•
1

100

•

' I

; i

' i

1 1

r i

) 1

0 /i

f L

? A?

400.000

300.000

mjooo

100.000

Fio. 17. — B.1C.F. waye-form plotted from the summed ordlnates of Fig. 16. AIbo the carre

of squared ordJnates.

Run a planimeter around the curve of the squared ordinates. Say that this
gives us the reading 2116. The square root of this is 46. Now run the plani-
meter around the dotted rectangle. Say the reading is 1346. We must multiply
this by 10, because we plotted to one-tenth scale. This gives us 13,460. The
square root is 116. Therefore the square root of mean square value of the E.M.F.
generated in twelve conductors by the full rectangular field-form, being taken at
116, the square root of mean square value of the E.M.F. generated in eight con-
ductors by the field-form shown in Fig. 16 is 46, and the ratio of 46: 116 is
0*396. This gives us Ke for a three-phase star-wound armature having a pole
of the dimensions given in Fig. 16.

To sum up, the process of finding K^ for a simple three-phase winding and for
any given field-form is as follows : Write down the values of twelve equidistant
ordinates which fairly represent the field-form, the maximum beinej tal«)n at 100.

THE MAGNETIC CIRCUIT 29

Take the sum of the first eight of these and go through the process shown on
page 27. Subtracting an ordinate from one end of the line, and adding the next
-ordinate and so on, obtain twelve summation values. These if plotted would give
the E.M.F. wave-form. Square each of these values, and plot to any convenient
scale. Kun a planimeter around the curve and take the square root of the reading.
On the same base line draw the rectangle with ordinate 1,440,000 plotted to one-
tenth scale. Run a planimeter around this rectangle, multiply the reading by 10
and take the square root The ratio of the roots is Kt-

If we go through this operation with the field-form given in Fig. 14, we obtain
the value 0*4 for K^. The same method would be employed for a two-phase
machine, but in this case we would take the summation of six ordinates instead of
•eight We will find that Ke for an ordinary two-phase generator with a field-form
like Fig. 14 will come out 0*325. It will be found that the above method gives
results sufficiently accurate, notwithstanding the fact that the number of slots per
pole is different from twelve. Where, however, there is only one slot per phase
per pole (a very rare case in modem machines), Ke would be multiplied by 1*045
for a star-connected armature and 1*15 for a mesh-connected three-phase armature.
Sometimes the arrangement of the conductors in the phases is not as simple as
shown in Fig. 16. There may be a short chord- winding as shown in Fig. 126, or
a single-phase machine may be wound with a band of conductors covering an arc
greater or less than two-thirds of the pole pitch. Whatever the arrangement of
the conductors may be, we can calculate the value of Ke by laying out the con-
ductors and the field-form as shown in Fig. 16, and after observing which of the
•conductors are in series with one another, and whether the ordinates of the flux
help or oppose one another in generating the E.M.F., making a summation at a
number of convenient relative positions of armature and field. It will be found
that it is convenient to take the ratio of the square root of mean square voltage
-actually generated in the conductors in series to the square root of mean square
voltage which would be generated in all the conductors arranged with a full pitch
winding moving in a uniform flux-density as great as the maximum flux-density in
the gap. The advantage of making our Ke a fraction of the hypothetical maximum
effect is that we see at a glance how much we are losing or gaining by any arrange-
ment of conductors, and we are less likely to make a mistake in the calculation
when we obtain a number whose reasonableness we can at once estimate.

We have seen that in those cases where the flux distribution curve is of sine
form, the value of Ke can be calculated at once without going through the process
<lescribed in the preceding pages, and the figures, 0*39 for the three-phase case
and 0*317 for the two-phase case, can be used as guides in guessing the value of
Ke when we have not time to work it out

In actual practice it is not necessary to go through the calculation like that
given above, except in special cases where the field-form is of a new shape. The
-constant Ke is known for the frame and for the type of winding we intend to
•employ. For common shapes of salient poles having the pole arc equal to 0*675
of the pole pitch and having the comers bevelled as shown in Fig. 16, the constant
Ke for a three-phase full-pitch star winding is 0*4, and this constant can generally
be used in rough calculations of all similar machines. The effect of deepening the

DYNAMO-ELECTRIC MACHINERY

bevel or of reducing the width of the pole is to reduce Kg, The effect of reducing
the bevel or of widening the pole is to increase Ke. We may take as a good
standard bevel one which is \ the width of the pole, and then work out the

•7 1

' \

'69

'68

'65

i
g

•64

■

•63

-Sfl

(

•6i

• /

///

•

•p/..

•6/

i /

■

'60

§.a

)

'57

'56

'54

1
t

'52

1 ,

'51

air

gap

atcc

irmer

-of Pi

ol&

'50

Bevel /

?/1 TIO ,
/3

air- gap at centre ofp6l&

iW i\s 1 /U

pole width

Fig. 18. — ^Values of £« for diffeient values of the ratio *~V' ".T u' and diffeient values of the

pole pitch

bevel ratio, the value of clg being 5. (See Fig. 10.)

values of the constant Kt for different depths of the bevel and different widths
of the pole. This has been done for a three-phase machine with a full-pitch
star-connected winding, and the results plotted in Fig. 18. From this figure
the values of Ke for various shapes of pole can be read off directly. The valuea

THE MAGNETIC CIRCUIT

given in the figure cover all the cases commonly met with in practice. For any
ease which is not directly covered it is easy to find a field-form falling under the
variables provided for under the figure, which has the same general shape and the
same area as the field-form of the case in question and whose shape is so nearly
the same that Ke will practically have the same value. The effect of chording the
winding is fully considered on page 113.

The effect of saturating the armature teeth is to make the field-form wider for
a given maximum flux-density in the gap, and this will affect the value of Ke. The
field-form is easily plotted by the methods described on pages 18 and 395. We
can then either square the ordinates and obtain the constant Ke as described on
page 27, or we can choose a field-form falling under Fig. 18 that has the same
general outline and the same area, and read off Ke with sufficient accuracy for all
practical purposes. The method of finding Ke for an induction motor is the same.
The shape of the field-form will depend on the amount of saturation, and if accuracy
is required the number of ampere-turns required for the teeth would be worked
out by the method considered on page 78. In general, the Ke for a full pitch
winding and with 25 % of the magnetizing ampere-tums thrown on the teeth
may be taken at 0*4 15. We will work out below the Ke for the induction motor
whose field- form is plotted in Fig. 15. From Fig. 15, by taking the means of the
ordinates of the two field-forms, we can get 12 ordinates which are proportional to
the following figures : 0, 128, 246, 325, 380, 405, 420, 405, 380, 325, 246, 128.

We can now go through the process described on page 27 with these ordinates
of the field-form, and thus obtain the ordinates of the E.M.F. wave. To make the
matter clear we give the figures below :

0
128
246
325
380
405
420
405

Squared
ordinates.

5,320,000

7-230

8350
8350

7230

5320

|2309|

subtract 405

1904
add -128

Squared
ordinates.

1 17761

subtract 420

1.356
add -246

3160

|2309|

subtract 0

2309
add 380

1 iiiol

subtract 405

705
add -325

1 380|

subtract 380

0
add -380

1230

[2689"]

subtract 128

2561
325

145

f2886l

8ubLract\^^T^-'

I 2886 J

1 380 1
subtract 325

-705
add -405

145

subtract . 325

2561
add 128

(2689]

1 -1110 1

subtract 246

-1356
add - 420

1 - 1776 1

1230

subtract 380

2309
add 0

|2b09|

3160

DYNAMO-ELECTRIC MACHINERY

The wave-fonn of the e.m.f. is plotted in Fig. 20. We now square the ordinates
of the E.M.F. wave-form and obtain the curve of squared ordinates. Whenever the
numbers become unwieldy we can divide by 10 or 100, because the actual values are
of no consequence. Now, if we had 12 conductors, each with an E.M.F. of 420, the
highest ordinate of the flux curve, we should have a maximum of 5060 instead of
2900. Squaring the 506 and plotting to a scale to bring it on the paper, we get

9Qjm

Fio. 20. — Wave-form of E.M.F. of induction motor having field-form as shown in Fig. 15, and

the squared-ordlnate curve of the same.

the dotted rectangle in Fig 20. We then run a planimeter around the squared
ordinate curve and then around the dotted rectangle, and remembering that the
rectangle is plotted to one-tenth scale, we find that the square root of the ratio
between the areas is 0*415. This, then, is the coefficient K^ for the induction motor
in question. If the saturation of the teeth had been higher, so as to give a field-
form with a flatter top, the value of Ke would have been higher. Values of 0*42 and
0*43 are not uncommon.

In oases where the wave-form is required more accurately than when found by the above
method, the designer may resort to the analytical methods which have been worked out in
ooncise form in a paper recently by Dr. S. P. Smith and R. S. H. Boulding.* With the
kind consent of these authors, an abstract of this paper is embodied here and on page 305.

It is shown on page 306 that where the flux is not pulsating the instantaneous voltage
generated in a full-pitch of Tc turns is

e=2rctrfBxlO-'* volts,

where v is the velocity, I is the length of core, and B^ is the flux-density at the position x
of the coil. Thus, the wave-form of the voltage in each conductor is the same as the wave-
form of the flux curve.

When the conductors of the armature phase-band are uniformly distributed (see page 305)

over an angle 2<r=-T (Fig. 321), extending at any instant between 6^ and O^y the mean

value of B throughout the coil span, 2<r, will be 5- / Bx(29. If now there are m coils,

*J<mm. Lh\E,, vol. 63, page 205 (1915).

THE MAGNETIC CIRCUIT 33

lettered a, &, c, ... to m in the phase-bond, and mTo=T total turns, the instantaneous voltage
in T turns will be

From page 22 we have Bx=BiSindx+B3 8in3^, + B5 8in5^x,

so that f *Bjrrf^= "I BjCos^x + xBjCosS^... etc. I *

= -|Bi{coe^2-co8^i)+ g5(cos3^2-co8 3^i) + ... >

= -2 ■! Bj sin -^^-J sm -^-J + -«- 8in3-2L_!8in3 ^*2~~^+--- r

6+0 6 — 0 •

Now ' ^ is the angular position of the centre of the phasd-band, and ' ' is equal

to half the angle subtended by the phase-band or coil breadth. We have denoted the angle
subtended by half the coil breadth by 0-, so that

,=»iZ»i=?| (Rg. 321); and let ^' = #.

Then y)" = 2 rrflO"" / B, 5H£ gin « + B, ?i^ sin 3« + etc. \

^^a y ff 3<r J

= 2 rW10-«{ Bi// sin ^ + Bs/a' sin 3^ + etc. } .
The coefficients such ab/^'^ — ^— - are the winding factors.

This expression shows us the efifect of spreading the winding. If <r=0, the wave-form
of the E.M.F. is the same as for B«, but as we widen the phase-band, making <r greater,
the values of the winding factors become smaller, since sin^0'<A<r, so that the higher
harmonics are reduced, and the wave-form of the E.M.F. becomes more sinusoidal.

The values of the winding factors for different widths of phase-band are given in the
table on page 307. Where the coefficients Bj, B2, B,, etc., are known, the wave-form
generated in a uniformly distributed winding can be readily calculated in the manner
indicated on page 308.

Where the conductors of the armature are not uniformly distributed (see page 305), but
lie in slots (there being a whole number of slots per pole), the expression for the instan-
taneous value of the sum of all the e.m.f.'s generated in the coils takes the form:

2"e=27V?n0-«{Bi/iSin^-l-Bj/38in3d+...-fB*/*sinn<?}.
But now the expression for the winding factors /i, /s, etc., is changed. Wo now have

sin Am -^

msinA^

where 7 is the angle subtended by one slot (see Fig. 324, page 310).

In this case /h does not always decrease as the order of the harmonic h increases, but
periodically rises to a maximum (numerically equal to/]) whenever k passes a multiple of 2Q,
Q being the whole number of slots per pole. This gives rise to ripples on the wave-form of
E.M.F. of the order h=2Q+l and 2Q-1, This is explained further on page 310.

Now the virtual value of the electromotive force,

where j^^y "^Sf ^^'f '"^ ^^® amplitudes of the several harmonics of the wave-form.

In three-phase star-connected machines £^=0 and £^=0, Where the fifth and seventh
harmonics are in evidence the graphical methods of determining iC* given in this chapter
will take care of them with sufficient accuracy. The harmonics due to the teeth are usually
of too high an order to have their effect accurately calculated by the graphical method,
but their amplitude is usually less than 5 per cent, of Ei, and we can see from the above
expression for E that the addition of a harmonic of 5 per cent, makes only a negligible
addition to the virtual value of the electromotive force, and could not be read on a voltmeter.
We therefore neglect the effects of high harmonics in determining the value of Jr«»

W.M. C

CHAPTER IV.

THE MATERIALS OF THE MAGNETIC CIRCUIT.

In this chapter and the next we shall deal with the magnetic properties of
iron and steel, and treat of the various parts of the magnetic circuit.

Following the course proposed on page 8, we shall as far as possible
employ the same methods in dealing with the magnetic circuits of all classes
of machines whether they be A.c. or c.c. generators, synchronous or asynchronous
motors. What we require are general rules for making calculations relating
to the air-gap, the teeth and slots, the armature iron behind the slots, the
pole limbs and the yoke.

The nnits employed. It is a little difficult to decide what units should be
employed in a book of this kind. Many designers in England and America
use inches for measuring the dimensions of their machines, and amongst these
some will employ kapp lines and others will employ g.g.s. lines for measuring
magnetic flux. Of these some will write "60,000 lines per square inch" and
others write " 60 kilolines per square inch." Some engineers, on the other hand,
prefer to make all their calculations in centimetres (using of course C.G.s. magnetic
units), and then where necessary to convert their centimetres into inches for
the British workman.

If the inch be taken as the unit of length, there is a great deal to say for
the kapp line as the unit of magnetic flux. The speed of machines is invariably
given in revolutions per minute, so that the formula,

volts X 10^ = revs, per min. x conductors x kapp lines x volt constant,

is very convenient, and in practice, the number of kapp lines being 6000 times
smaller than the number of C.6.S., is more convenient to write down and to
speak about than the number of C.G.S. lines. Thus one speaks of 10 kapp
lines in the gap and writes it down 10, instead of talking of 60,000 C.G.s. lines,
which must be written down either as 60,000 or as 60 kilolines.

All the above methods are so widely employed that we decided in the first
instance to illustrate the rules given in the book by working out one example
in each of the following systems of units :

(1) Dimensions in centimetres, magnetic flux in c.g.s. units.

(2) Dimensions in inches, magnetic flux in kapp lines.

(3) dimensions in inches, magnetic flux in C.G.s. units.

THE MATERIALS OF THE MAGNETIC CIRCUIT 35

This, however, was found to involve a great deal of repetition, and we have
therefore in the main employed the C.G.s. system of units, that being the system
which will probably be most generally employed in the future.

It is of course assumed that the reader is familiar with all the units with
which he is concerned in magnetic calculations, but we will give here for his
convenience a short statement of the relations between some of them.

UNITS OF MAGNETIC FLUX.

The unit magnetic flux, one C.G.s. line, has been named in America the
maxwell.

As one often deals in dynamos with many millions of lines, some engineers
prefer to work in Megalines, taking 1,000,000 lines as their unit Others take
Kilolines as their unit, and others again the volt-line or 100,000,000 C.G.S. lines.
The latter unit is very useful when speaking of the total flux of a frame. These
larger units, it is true, avoid the writing down of so many ciphers, and are
therefore useful in private calculation where the unit is familiar. In a book
on the subject, if one uses these units it is always necessary to write the word
mega or kilo in stating the units, so that much of the advantage is lost In
those calculations in this book in which we employ CG.S. units, we will use
the volt-line as the unit when dealing with the flux per pole or when speaking
of the total flux of a certain frame. We have, then,

100,000,000 maxwells = 1 volt-line.
1,000,000 maxwells =1 megaline.
1000 maxwells =1 kilolinie.

Dr. Gisbert Kapp in his early writings on the dynamo— writings with which
so many living designers are familiar — introduced the kapp line, which is
equal to 6000 C.G.S. lines. By its use we avoid the necessity of dividing the
revolutions per minute of a machine by 60 to convert to revolutions per second,
and we use the factor 10~* instead of 10"* in the well-known equati6n for the
voltage generated in a moving conductor. The kapp line being 6000 times
greater than the C.G.S. line, the number which expresses the quantity of flux
per pole in kapp lines generally runs to only three or four figures. At the
same time one digit is often sufficient to express the flux-density in the

gap-
Units of flax-density. A flux-density of one c.G.s. unit or one maxwell

per square centimetre has been named in America the gauss. In this book

we shall always write the flux-density expressed in C.G.S. lines per sq. cm.

as B. Where inch measurements of length are used it is convenient to write

B" for the flux-density expressed in C.GS. lines per sq. inch. If we have

occasion to employ kapp lines, we can write B^ for the kapp lines per sq.

inch.

We then have the following relations between these units :

1 C.G.S. line per sq. cm. = 1 gauss = 6 45 lines per sq. in.

36 DYNAMO-ELECTRIC MACHINERY

Or, as one is generally dealing with thousands of lines to the sq. cm., one gets
a better idea of the relation by writing:

10,000 CG.S. lines per sq. cm. = 64,500 lines per sq. inch.
10,000 CG.S. lines per sq. cm. = 10*75 kapp lines per sq. in.
10,000 CG.S. lines per sq. in. = 1550 lines per sq. cm.
10,000 CG.S. lines per sq. in. = 1*66 kapp lines per sq. in.
10 kapp lines per sq. in. = 60,000 CG.S. lines per sq. in.
* 10 kapp lines per sq. in. = 9310 CG.S. lines per sq. cm.

Units of magnetomotiye force. Most designers use the ampere-turn as their
unit of magnetomotive force and plot their magnetization curves accordingly.
The CG.S. unit is about 80% of this, it being necessary to multiply the

ampere-turns by — to convert into Jf, the magnetomotive force in CG.s. units.

We therefore have the following relations :

1 ampere-turn = 1*257 c.g.s. units of m.m.f.
1 CG.S. unit=sO"795 ampere-turn.

1 ampere-turn per centimetre on a uniform endless helix gives us a field
of intensity -3"= 1*257 inside the helix.

1 ampere-turn per inch on a uniform endless helix gives us ^=0*495.
If ^=1 inside the helix, the ampere-turns per inch = 2 02.

MAGNETIC PROPERTIES OF IRON AND STEEL.

The four chief materials with which the dynamo designer has to deal, in the
magnetic circuity are : Cast Iron, Cast Steel, Forged Iron and Steel, and Sheet
Steel.

Cast iron. Cast iron is used for yokes and spiders on account of its cheapness,
the ease with which it is cast into complicated shapes and the ease with which it is
machined. Though of much poorer magnetic quality than steel, it sometimes
pays to use a heavy section of it instead of a light section of steel. Sometimes it
happens that in big frames a great depth of material would in any case be necessary
in order to obtain sufficient mechanical stiffness and the magnetic qualities of cast
iron are then sufficiently good. For this reason cast iron is used to a great eictent
in the yokes of large continuous-current machines. Very often in slow speed A.c.
generators it is necessary to provide a certain amount of fly-wheel effect, and the
fly-wheel effect can be obtained most economically by employing deep cast-iron
rims on the field-magnet wheel. There being a great depth of material, the cast
iron is magnetically sufficiently good for the purpose. It is only at the root of
the poles that one feels the pinch, due to the poor magnetic quality of the iron.
Even where the number of ampere-turns on the magnetic circuit is somewhat
increased by the use of cast iron instead of cast steel, there may be cases where the
saving effected in using the cheaper iron pays for the cost of extra copper. Curve
6, Fig. 22, shows the relation between B and H for a fairly good specimen of grey
cast iron. Average cast iron, as commonly employed in dynamo frames, is not

THE MATERIALS OF THE MAGNETIC CIRCUIT

quite as good as this, if we take into account the whole casting including the
skin. One might take Curve 7 as an average curve ; if, however, the castings are
small and have been cooled quickly, the magnetic properties may be worse than

I i I I I I I I I i

T 1 1 1 1 \

"y9ut jLomts jt9d sawki

E
I

c
8

P
C
«
P

I I I I I i I I

i i i i i I s I I

T — I — I — I — I — I — I — I — ' — ' — ' — I — ^ — 1 — I~I — I — T — I — I — T — r~i

those shown on Curv^e 7. When cast iron cools, part of the carbon in it is deposited
in graphitic flakes, while the remainder is combined with the iron and has the
effect of greatly reducing its permeability. Very slow cooling results in a smaller
percentage of combined carbon. It thus comes about that two different pour-
ings from the same ladle may have considerably different magnetic properties.

38 DYNAMO-ELECTRIC MACHINERY

according to the way in which the iron is cooled. Good average grey cast iron
may have the following composition :

Graphitic carbon, - - - - 2*9 per cent.

Combined carbon, - - - - 03

Silicon, 2*5

Sulphur, 0-05

Phosphorus, 0-14

Manganese, 0*13

A sample of cast iron giving a curve as good as Curve 1 will have about 0*22
per cent, of combined carbon. The price (1914) of cast-iron dynamo frames in
large quantities, delivered in a Midland town, is from 9s. to 13s. per cwt. for
castings weighing between 1 and 20 cwt., depending on the difficulty of moulding,
and from £8 to £11 per ton for castings weighing between 1 and 10 tons. The
molten metal in the ladle may be taken at £6 per ton.

Malleable cast iron. When iron castings of no great thickness are heated to
redness for several weeks in the presence of haematite or manganese dioxide,
a considerable percentage of the carbon is burned out, and malleable iron is
obtained, possessing somewhat better mechanical and magnetic properties than an
ordinary cast iron. Curve 5 shows the magnetic properties of a sample of malle-
able cast iron. The quality of this material is, however, very uncertain, as much
depends upon the proportion of the combined carbon which has been burned out.
Malleable castings are conveniently used where it is necessary to have better
mechanical qualities than are found in plain cast iron, and where the pieces
are too small or too difficult to cast to justify the use of cast steel. They
are sometimes used for the end plates of poles, and for the finger-plates of
armatures.

A malleable casting having the permeability shown in Curve 5 might have the
following composition :

Graphitic carbon, - - - 2*0 per cent.

Combined carbon, . - . . 0*09

Silicon, ri

Sulphur, 0-01

Phosphorus, 0*03

Manganese, 0*08

The price of malleable castings (1914) depends largely upon the numbers
ordered, but may be taken roughly at 29s. per cwt. for reasonable quantities of
simple pieces weighing not less than 15 lbs. apiece. For smaller pieces the prices
may be higher, and will depend on the difficulty of moulding.

Cast steel. From its chemical composition one would expect cast steel to
possess very excellent magnetic qualities, and some samples of cast steel are as
good, magnetically, at the point of saturation ordinarily employed in electrical
machinery as forged steel; but, unfortunately, blow-holes and piping crevices
sometimes occur in the castings, and accidents may happen in the cooling which
bring about a rather poorer permeability. Curv^e 2, Fig. 22, shows the magnetic

THE MATERIALS OF THE MAGNETIC CIRCUIT 39

qualities after annealing, of a fairly good sound casting of dynamo steel, having
chemical composition as follows :

Combined carbon, - - - - 0*2 per cent.

Silicon, 0-15

Aluminium, 0-05

Phosphorus, 0*04

Sulphur, 0-03

Manganese, 0*11

The quantities of all these impurities, except the carbon and manganese, might
be doubled without appreciably altering the shape of the curve. An increase in
the percentage of carbon reduces the permeability. If the specimen had not been
annealed, the permeability at low inductions would have been lower, but the
permeability at about B= 18,000 would have hardly been affected. The addition
of manganese or chromium, or other hardening elements, reduces the permeability.
An addition of nickel up to 4 per cent, has no deleterious effect. Some steel
containing 2 per cent, of nickel shows a slightly higher induction at H = 100.
Experience shows, however, that one cannot rely upon always getting as good
material as is represented by Curve 2, and we may therefore take Curve 4 as the
curve of an average specimen of a steel casting. In this curve we have allowed
Sh per cent, of the space occupied, for blow-holes, and we have assumed that the
annealing will not be quite as good as in Curve 2.

There are several great advantages to be gained in the use of cast steel in
preference to cast iron. The permeability is so much higher that only one half of
the cross-section of material need be employed (assuming always that we have
sufficient mechanical stiffness), and the weight of the whole machine is greatly
reduced.

When the pole limbs are made of forged or rolled steel a smaller section of
limb can be employed where the yoke is of cast steel than if it is of cast iron,
because there is not the same fear of excessive saturation at the root of the
pole.

The cost of dynamo steel castings, delivered in a Midland town, is from 13s. to
15s. per cwt. for castings up to 10 cwt. depending on the difficulty of moulding
and the numbers ordered, and from £11 to £13 per ton for heavy yokes of simple
section.

It will be seen, in comparing these prices with the prices of cast iron, that if the
weight of the steel frame can be reduced to one half of the weight of a cast-iron
frame for the same machine, there is a considerable saving in the cost of material.
It is, however, usual to allow for more metal being taken off the finished faces.
Thus more cast steel goes to waste. The saving of freight on the completed
machine must also be taken into account.

Another advantage in the use of cast steel for dynamo frames lies in the fact
that the pole limb can in many cases be cast with the yoke, thus saving some cost
in machining. With cast iron it would be false economy to cast the pole limb
with the yoke, because the section of the pole limb if made of cast iron would have
to be made very large, and this would call for an excessive weight of copper for

DYNAMO-ELECTRIC MACHINERY

the winding. Steel castings have usually not as good a finish as iron castings,
and it is difiicult to make small sections and complicated shapes in cast steel.
The cost of machining of steel yokes is greater than the cost of machining cast^
iron yokes of the same size. Usually, with cast steel more chipping and prepara-
tion work is required. The cost of machining simple dynamo yokes of cast steel
and cast iron is shown in Fig. 23. The curves have been plotted from the records
of a large dynamo works where modem methods are employed. The figures for
cost include the cost of chipping and preparing for the boring mill. A certain

16
14

^ w

Cfi

200

1000

I2O0

400 600 800

Weight of yoke in lbs.

Fia. 22. — Cost of machining cast iron and cast steel yokes

1400

1600

noo

percentage (perhaps 30 or 40 per cent.) of the steel castings are defective and must
have the flaws welded. The average cost of this may be taken at about 3s. 6d.
per casting treated.

Another circumstance which must be taken into account in the choice of
material is the shortness of time in obtaining delivery from the maker. Steel
castings are only made by large steel manufacturers in certain centres, and it is
sometimes difficult to obtain delivery, whereas there is very little difficulty in
obtaining iron castings in any large manufacturing town, and most dynamo
builders make their own.

Forged steel and iron. Low carbon steel, when forged so as to make it
compact and homogeneous, is of all commercial materials the one to be relied

THE MATERIALS OF THE MAGNETIC CIRCUIT 41

upon for its magnetic qualities ; forged steel containing not more than 0*2 % of
carbon is practically as good as pure iron for the magnetic parts of generators.
Curve 7 gives the relation between B and H for a specimen of forged ingot iron
made by the open-hearth process, whose chemical composition is as follows :

Combined carbon, - - - - 0-15 per cent.

Silicon, ------ 0-06

Sulphur, 0-03

Phosphorus, 0*04

Manganese, 0*4

The specimen was annealed before testing. An unannealed specimen would
have shown a lower permeability at H = 10, but at H = 100 B would have been
practically as high as in Curve 1. The effect of adding carbon and other elements
which have a hardening effect is the same in forged steel as with cast steel. This
Curve 1 is about as good a magnetization curve as one can hope to get from any
commercial material. It is probable that there is no material which is 4 % better
at H = 100. A very pure specimen of iron thoroughly annealed would show
higher permeability at lower inductions, but the feature which helps the designer
in increasing the output of a frame is the permeability at fairly high inductions.
Some magnetization curves of steel that one sees have been plotted from
measurements which do not sufficiently eliminate errors, and the figures obtained,
particularly at high magnetizations, are often erroneous. The dynamo manu-
facturer, to be sure of his material, must test a specimen in an apparatus upon
which he can make a direct comparison with materials whose qualities he knows
to be good. Steel manufacturers' magnetization curves are useful as a guide,
but unless we know the method of measurement, and the individual who made
the test, too much reliance should not be placed upon them. In any case,
one cannot be sure that the specimen faithfully represents the bulk. The
only true test of the material of a dynamo yoke is a test of the finished
machine.

The main objection to the use of forged steel or iron in the construction of
dynamo frames is the cost of forging or machining the parts to the right shape.
The material in the rough is very cheap — rolled bars of rectangular or round
section can be bought at .£9 per ton delivered in a Midland town. Whenever we
can, without much labour, fashion parts of a dynamo, such as pole limbs, from the
rough bars, no cheaper or better material can be used. One is sure that the
material is solid, and one is fairly sure of the magnetic quality if the percentage of
carbon is low. The mechanical qualities are also good. For rotating field-magnets
which are to be subjected to very great centrifugal forces, it is possible to make a
steel containing not more than 0*4 % of carbon and 3*5 % of nickel, having as good
magnetic properties as are shown in Curve 2 in the higher reaches of that curve,
and possessing the following mechanical qualities:

Ultimate tensile strength, - - 45 tons per sq. in.

Elastic limit, - - - - 27

Extension of an 8 in. specimen, - 18 per cent

Reduction in area, - - - 40 „

7S.eoe

24^000

U.9O0

zaooo

17^000
fSjOOO

\ooe
a,9eo

Ampere-turns per Centimetre.

F:a. 23. — HagitetiutiOD curve o1

THE MATERIALS OF THE MAGNETIC CIRCUIT 43

If, however, the percentage of carbon does not exceed 0*2 (the nickel being
still 3*5 %), the ultimate tensile strength will be about 38 tons per sq. in. and the
elastic limit 20 tons. The magnetization curve may then be as good as Curve 1
in the upper reaches. For low values of H the permeability will greatly depend
upon the treatment which the material has had since the last annealing.

Sheet steel. The material from which dynamo sheet steel is rolled should be
very low in carbon. The following is the analysis of a good specimen of ordinary
dvnamo steel :

Combined carbon, - - - - 0*09 per cent.

Silicon, 0-01

Sulphur, 0042

Phosphorus, 0-089

Manganese, 0*36

The process of rolling it into sheets makes it if anything more compact and
homogeneous than forged steel, but at the same time a thin layer of oxide is
produced on the outside which, to a certain extent, reduces the permeability of an
iron core built up of sheet metal. Care should be taken that this layer of oxide
is not too thick. Sheet steel 0 06 in. thick, when reasonably clean and assembled
under pressures such as are ordinarily employed in the building up of pole pieces,
may be taken to be 95 % solid iron, the remaining 5 % is made up partly of oxide
and partly of air spaces between roughnesses of the surface. The dotted curve 3
in Fig. 21 may be taken as giving the magnetic properties of good average dynamo
sheet steel. An extension of this curve going up to very high flux-densities is
given in Fig. 23. If sheet steel 0-02" thick is papered with paper OOOIS" thick,
«uch as is used in the building of armature cores, the material can be compressed
under the ordinary pressure used in dynamo construction, until it has the per-
meability of material 92 % solid. Where the sheet steel is only 0-016" thick, and
is papered with the same paper, one cannot rely upon the solidity being more than
89 %, unless the steel is particularly clean and the pressure to which it is subjected
is very high. It is well for every manufacturer to make occasional tests of the
solidity of his built-up punchings, so that proper allowance can be made for the
space taken up by paper and air.

The cost of ordinary dynamo sheet steel may be taken at JBIO or £11 per ton.

Alloyed steeL In recent years a steel alloyed with silicon has come largely
into use for electrical machinery. The effect of adding between 1*8 and 5 per
cent, of silicon to an almost pure iron is to greatly increase its electrical resistance
and thus to reduce the loss in it due to eddy currents (see page 52). At the same
time this addition of silicon has a marked effect on the permeability and on the
hysteresis loss. The addition of silicon in a quantity less than 1 *8 per cent, seems
to slightly reduce the permeability of steel, but as the percentage is increased
from 1*8 up to 4*8 the permeability for inductions below 13,000 or 14,000 is
increased. This is shown in Fig. 24, which gives the magnetization curve * of two
specimens of silicon steel and also for comparison the curves of three other steels.

♦See
Alternate

Dr. S. Guggenheim, **The Ma^ietic Properties of Iron Alloys and their Uses in
3-current Design,*' The Electrician, vol. M, p. 539.

DYNAMO-ELECTRIC MACHINERY

one unannealed having comparatively poor permeability at low values of H^
another annealed cast steel and the third an annealed specimen of the softest iron
very low in carbon and silicon. It will be seen that though the specimens of
silicon steel have as much as 0*2 per cent, of carbon they are very permeable at
inductions below 1 3,000, and the greater the addition of silicon (below 5 %) the
greater the permeability at low inductions. At inductions over 14,000 the
addition of silicon lowers the permeability. This gives the magnetization curves
of silicon steel a very decided shoulder.

<fi^^

C.^9^

5*^
f*^

"^^

/ IP

V t

/ I It

1 ^' ,

/<>

]

•

nooo
i60oa\

MCOO

12000

lOOOC

9000

€000

4000

tooo

€0

100

Fio. 24. — Showing the high permeability of Bflicon steel at flux-densities below 18,000 and the

lower permeability at high flux-densities.

The effect of the addition of silicon on the hysteresis loss is shown in Table L
p. 48. The greatest loss at all inductions occurs in the steel alloyed with 0*18
per cent, of silicon. All these results were obtained from steels containing 0*2 per
cent, of carbon. Many of the alloyed steels on the market are very low in
carbon. A characteristic analysis of alloyed steel is as follows:

Carbon, 0*08 per cent.

Silicon, 3 0

Sulphur, 0-03

Phosphorus, 0*045

Manganese, 0*2

Although low in carbon, all these alloyed steels show a rather lower permea-
bility than ordinary steel at high inductions. This is a feature to be taken into
account when they are used in armatures with highly saturated teeth.

5»

>»

THE MATERIALS OF THE MAGNETIC CIRCUIT 45

One drawback to the use of silicon steel for making stampings is its great hard-
ness and brittleness. A steel containing as much as 4 % of silicon is very hard on
the dies, and sometimes the sheet breaks up under the punch just as hard cast iron
would. Even after the metal has been punched the teeth will sometimes break off.
Steels with a lower percentage of silicon are made by some of the makers, which
while preserving to a considerable extent the high resistance, and therefore the low
eddy-current loss, are at the same time easy to punch and perfectly safe under
bending stresses.

The hardening effect of the silicon, if it has not been carried too far, is of great
service in the armatures of high-speed machines.

The cost of silicon steel 0*5 mm. thick, having a loss under standard con-
ditions (see page 53) of 0*8 watt per lb., is from JB20 to JB23 per ton. For higher
qualities with losses as low as 0*56 watt per lb. the price ranges up to £S0 per ton.

LOSSES IN SHEET IRON.

The two losses occurring in iron subjected to an alternating-magnetic field are
{1) the hysteresis loss and (2) the eddy-current loss. When considering the
hysteresis loss a distinction must be drawn between an alternating field having a
fixed orientation in the iron and a rotating magnetic field, in which the orienta-
tion of the induction rotates continuously. The difference in the hysteresis loss
in these two cases has been investigated by Prof. F. G. Baily, and is clearly shown
in Fig. 25 reproduced from his memoir.*

At low fiux-densities and at flux-densities up to B= 15,000 the rotating field

^ves a rather greater loss than the alternating field, the general character of the

upward sloping curves being the same, but after we reach the value B= 16,000 the

losses produced by the rotating field decrease and come down almost to zero at

B = 20,000. The losses produced by the alternating field go on increasing up to

B = 24,000 after which they remain almost constant. The pure rotating field of

constant strength seldom occurs in practice. It would occur in the rotor of a

two-pole machine if it were not pierced by a shaft. In multipolar machines with

annular cores the orientation of the flux-density rotates as the machine revolves,

but the flux-density does not remain constant. The change that takes place

may be regarded as a rotation of the flux with an alternating flux superimposed.

The hysteresis loss, therefore, would be shown by a curve lying somewhere

between the two curves in Fig. 25. The relative strengths of the rotating

field and the alternating field differ at different depths in the core. In the

teeth we have chiefly an alternating field. As it would take too much time

in practical calculations to discriminate between the effects of the rotating and

the alternating fluxes, it is usual to compile curves (based on the actual losses

in machines) from which one can read off the number of watts per cubic cm.

or per cubic in. for a given flux^lensity and given frequency. Such curves are

given in Fig. 29.

*Phil^ Trans,, 1896, voK 187, A, pp. 715-746, "The Hysteresis of Iron and Steel in a
Rotating Magnetic Field."

DYNAMO-ELECTRIC MACHINERY

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Fia. 25. — Hysteresis loss in soft iron and liard steel subjected to aitemating and rotating fields.

► ■• •

THE MATERIALS OF THE MAGNETIC CIRCUIT

For an alternating field the hysteresis loss follows the well-known law of
Steinmetz, /F^ = ^B^"^, sufficiently well for practical purposes up to flux-densities
of 17,000 lines per sq. cm.

For higher flux-densities one must use a curve derived from experiment.
Fig. 26, when used in conjunction with the hysteretic constant, is useful in

0-9

•

0^5

/ ■

0-4-

•

Lithss per

sq. crru

n •'

5000 10000 15000 20000

B Lirvcs per sq. irv.

25000

50000

WOOOO

Bk Kapp Lines per sq. Ln.

I5O000

"T"

—I —
20

T y

0 ' ^5 h ' ' 15 20 25 30

Fie. 26. — Showing how the hysteiesis loss in iron increases with the flux-density.

giving the hysteresis loss up to any flux-density ordinarily employed in dynamos.
In this figure Kh is a function of B, such that Khxrj — the hysteresis loss in joules
per cycle.

Tahle I. (see p. 48) gives the value of the hysteretic constant for different
kinds of iron and steel.

Fig. 26 has been arranged .so that, whichever of. the three commonly used
systems of units is employed, the loss per cu. cm., the loss per cu. in. or the

DYNAMO-ELECTRIC MACHmERY

loss per lb. can be readily arrived at. The following are the constants to be
used in conjunction with K^ given in the figure:

Ajk X >y = joules per cu. cm. per cycle,
iffc X 7/ X n = watts per cu. cm. at frequency n,
16-4 xKhxrj X n = watts per cu. in. at frequency w.
b9x Khxrfxn — watts per lb. at frequency n.

Tablk I. Hystkretic Constants.

Material

Ilysterctic
constant =1).

Material

Hysteretic
constant =i|.

Good dynamo sheet steel

0-002

Silicon steel (Si) = 1-8%

0-004

Fair dynamo steel

0-003

The same steel (Si)=0'2%

00021

Silicon steel* (Si)=4-8%

0-00076

Very soft iron

0-002

— 4 V

0001

Cast iron

0-011 to 0-016

If* »» =3'o%

00013

Cast steel

0-003 to 0-012

>« i» —^ A

00016

Hardened cast steel

0-028

II II =2*5%

0 0022

Barrett's aluminium iron

0-00068

*See Dr. S. Guggenheim's paper referred to on page 43.

Example 4. What is the loss due to hysteresis in the armature iron behind the slots of a
25-cycle generator, the maximum flux-density in the iron being 11,000 lines per square cm.
and the volume of iron (which is of ordinary quality) being 250,000 cu. cm. ? From Fig. 26,
for B = 11,000 -fir* =0-29. We will teke the hysteretic constant as being O'OOS.

0-29 X 0-003 X 25 X 250,000=5400 watts.

Example 5. What is the hysteretic loss in the teeth of the same generator, the total
volume of the teeth being 1500 cu. in. and the average flux -density in the teeth being
140,000 lines per square inch?

From Fig. -26, /T* =0-955,

16-4 X 0-955 X 0-003 x 25 x 1500= 176 watts.

Example 6. Suppose that we were prepared to work the iron behind the slots of this
generator at 13^ kapp lines per sq. inch, how much extra loss would we have and how
many lbs. of iron would we save ?

250,000 cu. cm. of iron weigh 4320 lbs.,

11,000 lines per sq. cm. = 11*8 kapp lines per sq. in.,

— T— X yo:k=3770 lbs. giving a saving of 550 lbs.

From Fig. 26, for Bif = 13-5, A'jk = -36.

59 X 0-36 X 0-003 x 25 x 3770= 6000 watts.
6000 -54(X)= 600 watts extra loss at the higher flux-density.

eddy-current losses.

If we had a simple alternating magnetic flux through the sheet steel of an
armature, the direction of the flux being strictly parallel to the plane of the
laminations, and if the individual sheets were perfectly insulated from one
another, the eddy-current loss in watts per cubic centimetre of iron would be

;re = ^ X - X /2 X n2 X BS,„ X 10-i«,

(1)

THE MATERIALS OF THE MAGNETIC CIRCUIT 49

vliere p ie the specific resistancB of the iron, t the thickness of the sheet in

centimetrea, n the frequency and Bmu the maximum flux-density in lines per
sq. cm.

In practice, however, these conditions are seldom met with. The flux in

most dynamos partly alternates and partly rotates. The constant -^ = 1'645

should be increased considerably on account of this circumstance. Experiments
upon perfectly laminated iron, subjected to a magnetic flux changing as it does
in dynamos, indicate that the constant 2-8 is nearer the right value than 1645.
If we take the specific resistance of ordinary dynamo steel at the working
temperature (50*C.) as ITTxlO"*, we get the formula for the oddy-current
loss in watts per cu. cm. of iron,

»-, = 2-8 X ^^.~^-^^., x ;« x n^ x ^ x 10-"

= 2-4 X C X n" X B«,„ X 10-" (2)

We find, however, ttiat the measured iron loss in a completed machine is
usually much higher than the sum of the hysteresis and eddy-current losses calcu-
lated by the formulae given on pages 47 and 48. There are many reasons for
this. The sheet iron is often bent about after the annealing
in a w&y that increases the hysteresis loss. The insulation
between the sheets is by no means perfect. There may be
burrs on the edges which allow adjacent sheets to make
metallic contact, or the filing of the slots produces a similar
fifiect. It must be remembered that when the punchings
are assembled in a cast-iron frame the edges of the punch-
ings usually rest against the cast iron and make electrical
contact with it. If, therefore, through the filing of the
slote or from any other cause the punchings are in electrical
<M>ntact on the working face of the armature, there is a
complete electric circuit through which a current will pass /^f. jrari

driven by an electromotive force whose maximum value is ^^ 27 — Eddr-tmreDt n»th
«qual to 2vnN>i\Q"\ where N is the total flux carried
by the shoiircircuited punchings. Even if the punchings are insulated from the
frame by some thin, hard insulating material (a plan which may be adopted
with advantage when it is very important to keep down the iron loss), it is
possible to have a circuit along the burred punchings in front of a north pole,
with a return circuit along the burred punchings in front of a south pole. Or if
the two sides of a tooth are burred over, a current will flow around the electric
path thus formed, the electromotive force driving it being proportional to the flux
threading through that part of the tooth.

Another canae of excessiTe iron loss is the passage of the flux along a path
wbicb is not everywhere parallel to the plane of the laminations. At the ends of
the machine, part of the flux bulges out from the ends of the poles and enters the
armature on the flanks, and there is thus a considerable component of the flux at
right angles to the plane of lamination. This produces eddy currents both in the

DYXAMO-ELECTRIC MACHIXERY

end plates (whatever metal they are made of) and in the sheet iron. At the edges
of every ventilating duct the same sort of action occurs on a small scale. Again,
in armatures built up of segments there is always a little extra reluctance at those
parts of the magnetic path where the breaks in the punchings occur, even though
the punchings are arranged to break-joint. If from irregular machining of the
frame the punchings are built up so as to make a closer joint at one end of the
machine than at the other, as indicated in Fig. 28, the higher reluctance of
the joint at one end causes the flux in a certain measure to crowd to the end
where there is least reluctance. If now the bad joint is first at one end of the
machine and then at the other, there is a tendency for the flux to take a wavy

path, which necessarily has components at right angles to
the plane of the laminations. Sometimes the punchings
of the stator themselves build up so that the plane of
lamination is itself wavy, and the flux in each section of
the rotor as the machine rotates leaves and enters the
wavy stator punchings along paths which have small
components at right angles to the plane of those punchings,
and therefore causes some extra eddy-current loss. Wlier-
ever a break or partial break occurs in the punchings, some
of the flux is driven out into the surrounding frame, and
causes a little loss. It is well known that there are certain
relations between the number of breaks in the punching^
and number of poles which cause this loss to be greater or
less. If the number of poles is equal to number of breaks,
the relation is good. If the number of poles and the
number of breaks is such that there are at any instant
as many north poles opposite breaks as there are south poles opposite breaks,
the relation is good. If, however, the numbers are such that at one instant a
great number of north poles are opposite breaks (the south poles being between
breaks) and at another instant a great number of south poles are opposite breaks
(the north poles being then between breaks), the relation is not so good. Thus
we would not from choice have the number of poles 1^ times the number of
armature segments. This relation of the numbers is not impossible, but it is
to be avoided if possible, particularly if the joints in the punchings are not
well made.

Having regard to all the accidents that may happen, even in the best
regulated shops, we may be sure that the iron loss will be greater than the
amount calculated by the above formulae. It is therefore well to have curves
based upon actual experience from which one can arrive at the probable iron
loss. In these curves the hysteresis loss and eddy-cuiTent loss can be dealt with
together, and the total loss per cubic centimetre or per cubic inch can be read
off" directly.

The curves in Fig. 29 will be found useful for quickly estimating the iron loss
that may be expected in a built-up armature of ordinary manufacture. For the
purposes of these curves we have taken sheet iron 0*01 6"" (or 0*04 cm.) thick, papered
with 0*0013" paper, and assumed that the solidity is 89 %. The flux-density given

1' I/- .111 I'M!
■ III
' 1 ■

|l I'll

!i ■ ill

■

1 1 1

1 ' 1

!' ' 1 i 1
1 1 1

;■

FIO. 28.-
bnak-yyini
fluz-distril

1 1

—Showing mil
i which ajiecfa
lotion.

even
I the

THE MATERIALS OF THE MAGNETIC CIRCUIT

as abscissae is the actual flux-density in the iron. Thus the point 10,000 on the
abscissa refers to a state of magnetization in which we have a flux-density of
8900 lines per sq. cm. in the built-up mass of iron and paper or 10,000 lines

r-
0

VnO ^fiOQ ffiOO 9P00 HipOO a»00 mo m» moo ZqpOO 29p00 2^fi00 SepOO 29P00 9BtO0O

C6iS.Lmas per square Centimetre

-r*
4

6 lo 3 rs S S Ho S 55"
Capp Lines

Kapp

perstfuare inch/.

IS-

28 ^

Fio. 29. — Curves for quickly estimating the Iron loss in bailt-up stampings. Thickness of

stampings, 0'04 cm.

per sq. cm. in the actual iron. Good commercial armature iron has a hysteretic
constant as low as 0'0023, but after it has been punched and assembled we will
in general not find the constant much lower than 0'0027. The curves are therefore
based on this latter figure. In order to allow something for the short circuiting

52 DYNAMO-ELECTRIC MACHINERY

of punchings, which always occurs to a certain extent, we have taken the constant
3*7 instead of the constant 2*4 in formula (2), page 49. This constant gives
us a figure for the iron loss which agrees with the average case met with
in practice. For very carefully built-up armatures with very few short-circuited
punchings it is, of course, too high. On the other hand, many cases will be found
in practice where it is too low. The curves, then, have been plotted from the
formula,

watts per cu. cm. = (0-0027 xnxKk) + 3*7 (0*042 x w^ x Bj,„ x lO'ii),

and in this formula the values for Bm^x ^^^ the actual values of the flux-density
in the iron obtained by dividing the total flux by the net cross-section of the
iron.

It will be noticed that these curves have a curious knee in them, which occurs
near the point where B is about 18,000. This knee is produced by the fact that
the hysteresis loss does not increase much when we go above this density. The
knee is very marked in the curves for 15 and 25 cycles, because in these the eddy-
current loss is low as compared with the hysteresis loss. The curves as drawn
show us that at low frequencies we can go up to very great flux-densities without
being afraid of excessive losses.

It is, of course, impossible to give rules which will give the iron loss very
accurately. Two machines may be built to the same drawings and of the same
material so far as tests can show, and yet one may have 20 % more iron loss than
the other. In cases where a machine has received unfair treatment in punching
and building, its iron loss may even be doubled. The constants given above are
sufficiently near to obtain figures for the calculation of efiiciency. In cases where
it is necessary to give the very highest efficiencies, the actual hysteretic constant
of the material to be employed may be inserted instead of 0 0027, and the
coefficient 3*7 may be reduced to a value as near to 2 4 as is thought safe, having
regard to the amount of care that will be exercised in the building and treatment
of the core.

Where good silicon steel, containing 3 % of silicon, is employed, it is fairly safe
to take the hysteretic constant at 0*0016, but although the specific resistance of
the material is four or five times that of ordinary iron, it is not safe to reduce the
constant 3*7 to below 1*8 unless experience with similar armatures built with
the game care warrants it. Theoretically, with a perfectly built armature of
silicon iron, neglecting the losses which occur at the flanks (as to which a
separate allowance might be made), the iron loss per cu. cm. in an armature
might be reduced to

(0*001 X w X Kh) + 0*5(0*042 x n^ x 3% x lO'ii).

It is to be hoped that the day will come when our methods of treating and
building the iron will enable us to always use the last-given formula.

The ordinary method of stating the loss in any given sample of iron, is to give
a figure for the sum of the losses due to hysteresis and eddy currents in one pound
of the sheet iron when subjected to an alternating magnetic field with a maximum
flux-density of 10,000 lines per sq. cm. at a frequency of 50, the thickness of

THE MATERIALS OF THE MAGNETIC CIRCUIT 63

iron being 0*5 mm. The following list shows how the quality of the iron has been
improved during the last few years :

Materi&I.

Lo68 in watts per lb. under standard
conditions stated a)x)ve.

Dynamo steel in 1893,

Good ordinary (1914),

1-7

Better quality (1914),

1-3

Silicon steel (3 % Si),

0-9

Silicon steel (3*5% Si), -

0-8

Silicon steel (4-8% Si), -

0-56

The above losses are those which would be measured in the iron when built up
in a transformer core. For the reasons given on page 49, when the iron is built
up in a machine the losses are usually very much greater, as shown by the curves
in Fig. 29. It will be seen, for instance, that for the standard test conditions
(50 cycles B=10,0(K)) the loss given in Fig. 29 is 0*82 watt per cubic inch.
Taking the volume of 3*6 cu. in. of built-up punchings as weighing 1 lb., we have
2*9 watts per lb., or nearly double the figure given above for good ordinary iron.
One recognizes, therefore, how important it is to build the iron carefully and keep
it free from burrs. The curves given in Fig. 29 are average curves ; the losses can
be easily exceeded on a badly burred core.

It is the practice of some manufacturers to anneal the sheet iron after punching.
This has the effect of preventing the increase of hysteresis loss which may have
been occasioned by the straining of the metal under the punch, and it also has the
good effect of. oxidizing sharp edges burred up by the punch. Very low iron
losses have been obtained with sheet metal so annealed, even when the insulation
between sheets consisted only of varnish. The objection to using only varnish
between the sheets is that the varnish may in the course of time be squeezed out,
so that burrs and projections on the punchings make contact with one another,
and the iron loss is thereby increased. A solid insulator like paper makes a more
permanent spacer between sheets.

It is convenient to paste the paper on the sheet before it is punched, and one
cannot anneal papered sbeet iron after punching. It is found that if sheet metal
is not annealed after punching it builds up more accurately. The punchings are
sometimes slightly distorted during the annealing process in a way whith causes
them to build up less accurately, and this gives a rougher surface inside the slots.

Silicon steel is frequently rolled to sheets that are rather thicker than the trans-
former iron of the ordinary sort. The specific resistance is from three to five
times as great as ordinary iron, so it is not worth while to roll it so thin. A
thickness of 0*5 mm. is common. If the percentage of silicon is increased to 4*8,
the resistance goes up to six times that of ordinary iron. Such a high percentage
of silicon, however, makes the steel too brittle for use in dynamos.

The curves given in Fig. 30' may be taken as giving average losses per cubic
inch alloyed sheet steel (percentage of silicon, 3 per cent.) assembled in ah
ordinary armature core and subjected to fair treatment.' Alloyed sheet steel
containing 4 per cent, of silicon is sometimes rather brittle, and is rather difficult

DYNAMO-ELECTRIC MACHINERY

to punch. Moreover, the punchings may become brittle with time even if they are
not when the sheet is punched, so that the teeth break off when subjected to
bending forces. This is a very dangerous fault. In order to meet this difficulty

SO
4^

44
42
4-0
38
3-6
3^4
3-2

3i>
Z8

t6
[24-

1/^

•s

5? /•♦

'029

ir£0

-0-27

-0-2B

-0-25

-0-24
-0-23
-0-22

-0-2I

-0-2D

-0-19

-0-m

'0-17
'O-W

■^

HW5

-OH^

•s

-iH3

.nJv

•

-o-n

CT//

-Oi8

•&C9

-Mf

iHfT
HHf6
•4Hf5

-004

-iHf3

-iHfZ

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spe

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inci

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0-1

O 2 4- 6 d W 12 14- W 18 20 2Z 24 26 28 30 32
0 2000 4000 6(Hn 8000 10000 080$ 14^

B Lines per sq. centimetre

Fio. 80. — Cnrves for aniokly estlmattng the iron loss oociurlng In ailioon steel (8 per oent.
flUloon) oHflonibledln an ordinary armatnre core. Thicknees of stampingB, 0*05 cm,

■

some steel manufacturers make an alloyed steel with a rather smaller percentage
of silicon (about 3 per cent). This material, though not at all brittle, has a tensile
strength as high as 105,000 lbs. per sq. in., and is therefore very suitable for the
rotating armatures of turbo machines.

For references to articles on dynamo steel and iron loss, see page 86.

CHAPTER V,

THE PABTS OF THE MAGNETIC CIRCUIT,

In this chapter Tve will collect the rules which are of service in making calcu-
lations relating to various parts of the magnetic circuit. The following symbols

Fio. 81. — Parts of the magnetic circnit.

will be used in this book to denote the lengths of the various parts of this

«^^^^*- ^ = length of gap,

Z2 = length of teeth,
Za. = length in armature core,
2p = length of pole body,
/2^ = length in yoke.

Fig. 31 gives a sectional view for the magnetic circuit of a revolving field A.c.
^nerator.

56 DYNAMO-ELECTEIC MACHINERY

THE AIB-GAP.

The flux-density in the air-gap. We have seen in Chapter II. how we can plot
the flux-density along the pole face in various types of machines. It is usual, in
calculating the electromotive force generated in the conductors of an armature, to
regard the conductors as moving across a magnetic field in the air-gap of the
machine. The results thus arrived at are in the main correct, although the con-
ductors may not be in the air-gap but in slots. The total flux cut per pole is
the same whether the conductor is actually in the strong field, in the air-gap, or
in a weak field in the slot. We may satisfy our notions of a conductor-moving-in-
a-field by saying that the velocity of the weak field in the slot is greater than the
velocity of the periphery of the armature in the inverse ratio of the density in the
slot to the density in the gap.

Thus, our formula E = KeBAgZtRpm^ ^j^ x^ 10~^ given on page 7, holds for all
machines, whether surface wound or iron clad, and whether the conductors are
mechanically driven through the field or the field rotates magnetically as in an
induction motor (see p. 304).

The fiux-density in the air-gap of a machine may be taken as a convenient
criterion of the good use that is being made of the magnetic circuit, and for any
given frame its value tells us of the state of saturation of the machine.

Thus, if the flux-density in the air-gap of a certain machine is 30 kilolines per
sq. in., and if we find that at full speed we generate 3 volts per conductor and
have 60 kilolines per sq. in. in the teeth, 40 in the iron behind slots and 45 in the
pole limbs, then with 60 kilolines in the gap we shall generate 6 volts per conductor,
and have as a first approximation 120 kilolines in the teeth, 80 in the iron behind
slots and 90 in the pole limbs. The gap-density is a convenient quantity to which
we can refer the intensity of the magnetic effects in all parts of the machine,
although in many cases account must be taken of leakage fluxes, which interfere
with a true proportionality between the various quantities.

Its value, moreover, tells us at a glance whether a given frame is being used to
its best advantage. We know that in many cases the flux-density in the gap may
be as high as 60 kilolines per sq. in. at no load. If the figure is lower than this
we will not be satisfied with the design until we have found a sufficient reason for
having it so low. Or we may have it above 60, in which case we may get a
proportionately higher output from the frame.

What, then, is it that limits the value we may take for the flux-density in the
gap 1 Most commonly it is the excessive saturation that would occur in the iron
of other parts of the magnetic circuit if the air-gap density were too great But
this is not always the limiting condition. In some large alternators with a great
number of poles and a small air-gap, the flux-density in the gap cannot be increased
beyond, say, 50 without making excessive the unbalanced magnetic pull for small
displacements of the frame from the true concentric position. In this case the
output of the frame may be limited by this consideration. With induction motors
too great a flux-density in the gap would call for too great a magnetizing current,
and with wound motors the density must sometimes be kept low to prevent an
excessive magnetic pull. These matters will be dealt with in their place, and in

THE MAGNETIC CIRCUIT 67

the designs worked out in the subsequent chapters the reader will see what
consideration it is that limits the value of the flux-density in each particular case.

As the possibility of an unbalanced magnetic pull must be considered in both
continuous-current and alternating-current generators and motors, we will deal
with it here.

Unbalanced magnetic pull. As long as the armature of a generator or motor
remains concentric with the field and the frame does not become distorted,
the poles exert an even magnetic pull up and down, right and left, for
each carries the same number of ampere-turns. As the upward forces are
balanced by the downward forces, the bending moment in the shaft is produced
only by the weight of the rotating part. But this is a state of affairs that we
cannot always count upon. The bearings may wear and let the rotor down a small
fraction of an inch. Some small initial dissymmetry may bring about the springing
of the frame, and as the air-gap closes up on one side the magnetic pull there may
increase at such a rate that it is able to pull the armature hard up against the
field-magnet. Sometimes a dissymmetry in the winding or in the quality of the
material is sufficient to start the trouble. It is therefore necessary to calculate
how much the unbalanced pull amounts to when we have a small accidental
displacement, and make such provision in tlfe design of the shaft and frame as will
with certainty {Hrevent a pull-over.

A simple plan is to assume that if the shaft and frame are strong enough to
withstand the unbalanced pull which would be caused by a displacement of, say,
1 mm. if we are using G.G.s. units (or, say, -^^ inch) from the true concentric
position, then it will be strong enough to withstand the accidents of this kind
which may happen in service. The assumption enables us to give to the designer
of the mechanical parts a definite figure for the magnetic pull, and this figure he
adds to the weight and other forces on the parts when calculating the maximum
deflexion. This deflexion must in general be well within the 1 mm. or ^^ inch
as the case may be. We must not, however, forget that special cases may arise
in which it is necessary to make provision against displacements greater than 1 mm.
The unbalanced magnetic pnll dae to a small displacement of the armature.
We can deduce by the method given below a convenient formula for calculating
the unbalanced magnetic pull.

It should be pointed out, in the first place, that the amount of the unbalanced
pull for a given amount of displacement will depend upon the extent to which
the iron parts are .saturated. If the iron parts of the magnetic circuit are very
much saturated, a reduction of the air-gap on one side of the armature will not
result in a very great increase in the flux-density in the gap. If, however,,
there is no saturation, then the flux-density will increase in inverse proportion
as the gap is shortened. Other things being equal, the unbalanced pull will be
greatest for an unsaturated magnetic circuit. This is the easiest case to calculate,
so we will take it first. It is then possible by a simple approximation to allow
for the diminution of the pull due to the fact that some of the ampere-turns
are expended on the iron.

From first -principles we know that the pull on the face of a magnet (made
of a material of~"great permeability), per square centimetre of active face, is.

^8 DYNAMO-ELECTRIC MACHINERY

o- dynes. This is easily seen when we remember that the energy stored in a

1 HB
•cubic centimetre of air-gap is -x -r— ergs (see Elements of EledricUy and Magnetism^

by J. J. Thomson, 1893, p. 266). This may be put into the form ~— or -tt-.

As ft=l, in air, the energy = g-. Now imagine that the magnetic pull makes

the iron move so that the space that was air-gap becomes occupied by the iron

K)f great permeability, fi. If B remains constant H becomes nearly zero, so that

B'
the energy stored per cubic centimetre becomes -— >, that is to say, nearly zero.

In order to thus convert the magnetic energy in one centimetre cube of air
into mechanical work, it is necessary to move the square centimetre of iron
surface through the centimetre, so that the force exerted must be

l(l-l) dynes.

It is only in the case where the permeability is great that we can neglect
-the term -. This is not the case when the iron is very highly saturated.

The pull in lbs. per sq. in. = Q ^^ — 7»« = 5'75x 10"'x B*.

Or, if the flux-density is measured in lines per sq. in.,
the pull in lbs. per sq. in. = l-39 x IQ-^x (B")*.

If the flux-density is given in kapp lines per sq. in.,

the pull in lbs. per sq. in. = 1-39 x lO'S x (6000)* x B

The last expression is in such a simple form that we will keep to the Bk

\inits in what follows.

Let Mp equal the magnetic potential between the pole and the armature

Mp
of a dynamo, the units being so chosen that — = B^ , where g is the length

-of the gap between the pole and armature in inches.

B^ M^
Now the pull in lbs. per sq. in. = -^«=^«

Consider the difference in pull at two diametrically opposite points at which
-the gap is (g-a) and (g-^a) respectively (see Fig. 32),

Diff'erence in pull= -^''(^-J-^- -^_L_) ibe. per sq. in.

Consider first the case where a is small compared with g.

Expanding the expression and neglecting quantities much smaller than ^,

the difference in pull = -^ x --3.

But M= Bjc X g, therefore the difference in pull per sq. in. = -^ x -, = — ^ •

2 g^ g

THE MAGNETIC CIRCUIT

Referring now to Pig. 32, as 6 changes, the distance between the dotted circle
and the full circle changes as a sin ^, and the vertical component of the difference
in pull varies as asin^^.

Let the area of any pole surface subtended by the angle d be equal to ^^.

Fio. 32. — Diagram of fldd-magnet displaced from central podtioD.

Then the total difference in pull taken half-way round the circle

Jo g

2B%aA

9 Jo g

VT ^ . i_ 1^ .t X 1 1 e No. of poles xsq. in. of pole area
Now At IS half the total polar surface = -^ .

The difference of pull in lbs, = 0*5 x — ? x No. of poles x sq. in, of pole area.

= 0*5 X B^ X No. of poles x sq. in. of pole x ->
where a is very small* as compared with g.

Example 7. A 300 k.w. d-phase generator has 12 poles. Each pole face has an area of
70 square inches. The length of the air-gap is 0*18* and the normal flux-density in the gap is
58,200 lines per sq. in. Find the unbalanced magnetic pull when the field is displaced -^'^
assuming that there is no saturation of the iron.

58,200 lines per sq. in. ^9*6 kapp lines per sq. in.

Unbalanced puU in lbs. = OS x 9-6 x 96 x 12 x 70 x -^^

=6760 lbs.

'^Or, to be more accurate for saps of 0*1 inch and under, where - is not necessarily very

11 g

snialL we should preserve the expression r^-- r^.

ig-a)^ ig+a)^

Then the difference in pull in lbs.

= 0*125 B,flr^ { ;a — h\ X No. of poles x sq. in, of pole area.

\to-ar {g+afj

If ^=the number of thirty-seconds of an inch in the gap, then, for a displacement of one
thirty -second of an inch from the central position, the unbalanced pull in lbs.

=0*5 B^ X Na of poles x sq. in, of pole area x I t +^ ]•

60 DYNAMO-ELECTRIC MACHINERY

If the flux-density is given in lines per square inch, we must change the
constant 0*5 to 1-39x10-8.

Or, if the fiuzden8ity is given in lines per sq. cm. and the area of the pole
in sq. cms., the magnetic pull in kilograms is

4*05 X 10~® X B^ X No. of poles x sq. cm. of pole x -•

Example 8. A certain 3-phaae generator driven by a gas-engine has 60 poles each having an
area of 450 sq. cms. We want to have about 5000 ampere-turns on the pole at no load on the
air-gap, and it is required to keep the unbalanced magnetic pull due to a displacement of 0*1 cm.
less than 13,000 kilograms. What is the maximum flux-density in the gap that we can employ
and what will be the approximate length of gap, assuming no saturation of the iron ?

w 1 • *u fi * 1 5000x1-267
We have, in the first place, g= g

Therefore 13000=4-05 x lO"" x B' x 60 x 450 x —MiL^^.,

5000 X 1*25/

6=9100, (7=0-69 cm.

The effect of saturation on unbalanced pull. In the above we have assumed
that all the ampere-turns are expended for the air-gap. Now, let us see what
modification is necessary where a considerable percentage of the ampere-turns are
expended on the iron parts of the circuit.

In the case of very large slow-speed generators, the number of poles being
great and the economical air-gap small, the unbalanced magnetic pull would often
be almost too great to cope with if it were not for the fact that the actual pull is
very much less than the pull calculated by the simple formula given above. We
cannot, in these cases, neglect the effect of saturation. A simple graphic con-
struction enables us to allow for its effect with sufficient accuracy for practical
purposes. This construction is based on the argument that whatever the amount
of the unbalanced pull may be under the effect of saturation, we can always
imagine an air-gap of such a size that the unbalanced pull would be the same,
with the same flux-density and no saturation. The object of the graphical
method is to find out what the length of this equivalent gap is. The method is
most easily understood from an example worked out

Fig. 347 gives the no-load magnetization curve of the 1800 K.V.A. three-phase
generator, particulars of which are given on p. 357. The length of the air-gap
(i total gap) is 051 cm. The flux-density at 6600 volts no load is 9160 C.G.s.
lines per sq. cm. It is required to find out what the unbalanced pull will be when
the centre of the rotor is displaced by 0*1 cm. from the central position.

First find how many ampere-turns are required to drive the flux in the gap
across 0*1 cm. We have

0-8 X 9160 X 0-1 = 733 ampere-turns.

Draw the two small vertical lines, as shown at c and d, near the working part of
the magnetization curve. Take as the horizontal distance between these lines the
distance on the horizontal scale that represents 733 ampere-turns. Now draw a
chord to the curve through the two points intersected by these lines, and through
the origin draw the line Ob parallel to this chord. Let this line Ob intersect at
the point b the horizontal line ab, drawn at a height to represent the working

THE MAGNETIC CIRCUIT 61

f)ux-density or voltage. Then the ratio of the line cd to the line ab is the ratio of
0*1 cm. to the length of the equivalent gap. For it is easy to see that if the gap
were so great as to require 11,900 ampere-turns at no load, the increase of the flux
for a decrease of the gap of 0-1 cm. would be the same as the increase of the flux
in the actual machine for a decrease of the gap of 0*1 cm. We can therefore

-employ the same formula as before, except that instead of the ratio - we use the

a .

ratio — , where g^ is the length of the equivalent gap as defined above. In our

example we get ~ from the ratios of the ampere-turns rr|-^7r7: = 0061, so that our
formula gives us

4-05 X 10-8 X 9160 X 9160 x 40 x 650 x 0-061 = 5400 kilograms.
If we had neglected the eflfect of saturation, otlr formula would have given us

5400 X ^-'1 , = 1 7,700 kilograms.

Permissible amount of unbalanced magnetic pnlL In very large engine-type
alternators the designer of the mechanical parts should be provided with a curve
showing how the magnetic pull varies as the displacement varies from zero to
(say) 0-1 in.

In order to give a general idea as to how great a magnetic pull is permissible,
we may say that the unbalanced magnetic pull with ^^" displacement in the case
of engine-driven alternators of from 100 to 500 K.W. capacity may be as great
as the weight of the rotating part, but to have it as high as this necessitates the
use of a rather strong shaft. Usually the unbalanced pull does not amount to
80 much.

Effect of circtdts in parallel on unbalanced pull. In the case of c.c. armatures
with a number of circuits in parallel, particularly where a great number of equalizing
<K)nnection6 are employed, there cannot be any great magnetic pull when the machine
is rotating, because if the field were stronger on one side than on the other it would
set up currents in the windings of the armature which would tend to weaken the
field on the side where the small air-gap tends to make it strong and strengthen it
on the side where it would otherwise be weak. It is therefore usual to neglect
the unbalanced magnetic pull in such machines. But it must be remembered that
«uch armatures may be subjected to an unbalanced pull when stationary if they
are separately excited. Similarly, all armatures, whether for A.c. generators or
induction motors, having several circuits in parallel, or having a short-circuited
winding as a squirrel cage or amortisseur on the rotating element, cannot have a
much stronger field threading through one part of the circuit than threads through
the other. Sometimes alternator armatures are wound with a number of circuits
in parallel with the express object of neutralizing the unbalanced magnetic pull
(see page 452).

It is not sufficient to make several circuits in parallel on the rotor of an
induction motor if we wish to obviate the unbalanced pull, because the slip may
be so small that the resistance of the rotor circuits will not permit a sufficient
•equalizing current to flow.

62 DYNAMO-ELECTRIC MACHINERY

Length of air-gap. In the calculations of the different machines given ii^
the subsequent chapters, the reasons for iixing the length of air-gap at the
chosen value will be made clear. It is only necessary here to briefly state the
various considerations which influence the designer in settling upon the length
of air-gap.

(1) To get sufficient magnetomotive force on the field-magnet. In A.C. generators and
c.c. generators and motors without compensating windings, it is necessary to have
the magnetomotive force of the field coils much greater than the magnetomotive
force of the armature, in order to secure good regulation and good commutation.
For this reason the air-gap is often made very much greater than it otherwise
would be. For instance, one sometimes sees air-gaps of 2 or 3 inches in turbo-
generators which have very few poles.

(2) To keep down the iron loss. In machines with open slots, either on the rotor
or stator, the air-gap must not be reduced below a certain minimum or the iron
loss will be excessive. This must be taken into account not only in small c.c.
machines with solid poles, but in all machines, whether the poles are laminated or
not. Where any iron is made to pass in front of a number of magnetized iron
teeth with only a short air-gap intervening, there is a change in the flux-density
which has the frequency of the passage of the teeth, and this is usually much
higher than the frequency of the passage of the poles, and gives rise to excessive
iron losses.

(3) Mechanical considerations. The air-gap must always be large enough to
obviate any danger of the rotating part coming in contact with the stationary
part, either from the wearing of the bearings or from the springing of the shaft or
frame under the action of magnetic pull or other forces to which the machine is-
subjected. The air-gap in railway motors and in induction motors is fixed by thia
consideration.

(4) T'o ventilate and obviate noise. In some machines, such as turbo-generators,
in which a great deal of air must pass along the gap, the minimum length is some-
times determined by this consideration. Where the pole faces of a machine are
large and occupy a great proportion of the space around the armature, the air from
the ventilating ducts in the core would not find a sufficiently easy path if the
air-gap were made as small as perhaps it might be made from other considerations.
Even when numerous air-ducts are provided in the stator core, the air-gap must not
be made too small or the blowing of the air against small projections on the stator
or rotor will cause excessive noise.

(6) I'o keep down tJie value of the unbalanced magnetic pvM, We have seen above
(page 59) that the unbalanced magnetic pull for a small displacement from the
concentric position is inversely proportional to the length of the air gap. In very
large engine-driven alternators in which the magnetic pull is a determining factor
in the design, special consideration must be given to the length of gap and ita
influence on the amount of the pull. Where a small fiux-density is employed in
the gap, a greater gap can be employed with a given number of ampere-turns oa
the pole (see page 347).

The ampere-turns on the air-gap. The simplest case to consider is where the
face of the armature is smooth and the pole is free from ventilating ducts and

THE MAGNETIC CIRCUIT 6*

slots. The ampere-tums required to create a flux-density of B IIdob per square
centimetre across a gap of g centimetres is

i^xBs- or 0-795BJ7.

Or, if we are working with dimensions in inches and with B' measured in lines
per square inch, then

ampere-turns on the gap = 0313 x B'x/,

ExAVPLl 9. Id a certain generator the value of OTA, was 300 megoliDes. Taking
Ag as 4900 eq. in., we get B'=61,000 lines per aq. in. If the gap is 0-376 inoh, then

ampere-tums on the gap^O-SlSx 61.000 xO'.t7S = T200.
Or, again, if n-e wish to work in kapp lines and have B;i kapp lines per square inoh, then
ampere-tums on the gap = lSSOx Bjr xg".

It will he seen that method of calculation described ou page 7 ohviatea all
the calculations as to the number of teeth per pole or the area of the pole, or the
density of the flux under different parts of the pole. It takcB into account only
the maximum density in the gap, and leaves to the constant K, the duty of making
allowances for the width of the pole, the bevel of the pole and other matters which
affect the total electromotive force generated.

It should be remembered that very often the calculation given above is reversed
in practice. It is often necessary, say for the purpose of securing good regulation,
to apply a given number of ampere-tums to the pole. Id this case we make the
air-gap great enough to call for the desired number of ampere-tums.

BxAHPLE 10. In a certain a.c. generator it is desired to have 10,000 ampere-tums per
pole. It is also desired to throw 10% of the ampere-tums at no load on the teeth for the
purpose of getting the desired saturation. Deducting 10 % from the 10,000, we have 0000 for
the gap. Assume that the flux-density is 61,000, as in the last example,
9000 = 0-313x61,000x3",
3* = 0-47 in.
The efltot of open slots and Tentllating dncts. Next consider the case where
there are open slots in the armature. The effect of the open slots is to increase

Fia, $9. — Showlikg bow m*sncUc Buz bom unutun teeth dlstributa Itself [n the air-gap.

the reluctance of the air-gap. The lines of force crowd into the tops of the toeth
in the manner indicated in Fig. 33. The effect of the slots in thus contracting the

•64

DYNAMO-ELECTRIC MACHINERY

path of the flux across the gap depends upon the ratio between the width of the
«lot and the length of the gap.*^ We may calculate the amount of the contraction

ui:

rrrr

III l'

iiil

^^-OS

I
I
I

^\i

i^ — as

■\ s .

TTTT

,M' I

TfF

A^JLL

I
I

lillL_i

Fio. 84. — Explaining the convention upon which the curves in Figs. 36 uid 37 are based.

of the path by considering that, for a region somewhat narrower than the slot, no
flux passes at all, and that for the remainder of the pitch of the slots the flux is
uniform, as shown in Fig. 34. The values of the coefficient o-, by which we must

TTTT

n n

width cfslot ,
length of gap "^

Fig. 35.7— Curves giving the values of <r for various ratios of width of slot to length of gap.

multiply s in order to get the virtual value of the slot width on this assumption,
are given in Fig. 35 for different ratios of slot width to gap length.

• See papers by F. W. Carter, Jctamal of Inst. Elec. EnffineerHy vol. 29, p. 925, and vol. 34,
p. 47 ; also Elec, World cmd Engr., Nov. 30, 1901 ; Hawkins and Wrightnian, ibid,, voL 29,
p. 436 ; Hele-Shaw, Hay and Powell, ibid,^ vol. 34, p. 21.

THE MAGNETIC CIRCUIT 66

The full-line curve gives the value of a- for rectangular open slots. The dotted
curve gives the value of u for semi-closed slots of the kind usually met with in
practice. Strictly speaking, the value of a- depends upon the shape of the lips of
the teeth, but for practical purposes the two cun'es given are sufficient.

Taking the value so- and deducting it from the slot pitchy,, we get the effective
tooth width shown in Fig. 34. On the assumptions made in Fig. 34, the flux-
density in the air-gap is increased by the presence of the slot, so that instead of

being B it becomes B x -^^ . It is convenient to have curves, such as those

^ven in Fig. 36, from which we can take the contraction coefBlcient Kg without

calculation. This curve is used in the following way : Suppose that in the

example given on page 23 B" is 60,000, the length of the gap 0*2", the width of

0'4
the slot 0-4'' and the pitch of the slot 0*8". We find the abscissa 7^:^ = 2. From

s 0*4
2 we run up the perpendicular until we come to the curve - = — = 0*5. The ordinate

of this curve where it meets the perpendicular 2 is 1*16. Therefore 60,000 x 116
is the apparent flux-density in the gap, so that the ampere-turns in the gap under
the above conditions would be

0-313 X 60,000 X M6 x 0-2'' = 4360.

The ampere-turns are 16 per cent, higher than they would be if there were no
open slots. If there are ventilating ducts in either the rotor or the stator, or both,
we can find the apparent contraction in the flux path caused by them in the same
way, as will be seen from the following. Let the gross length of core in the above
example be lO"", and let there be four ventilating ducts, each Y ^d^* Here we

have _ — — rr — J = -7r-7r-=l '26. The sum of the widths of the ducts is T, and

length of gap 02

as these are spaced over 10", we may take - = t7^' Now take the curve y\yth

where it cuts the perpendicular from 1-25. This gives us the ordinate 1*02, the
contraction ratio due to the effect of ducts, so the ampere-turns become

0-313 X 60,000 X M 6 X 1-02 X 0'2'' = 4450.

Similarly, if there are ducts in the stator as well as in the rotor, the contrac-
tion ratio for these is found in the same way. The usual practice is to find
separately the contraction ratios for the rotor slots and ducts and the stator slots
and ducts respectively, then the product of all four gives the total contraction
ratio. After a little experience with Fig. 36, an allowance for the smaller con-
traction ratios due to the ducts can be estimated correctly enough without
referring to the curve. The designer allows 2 or 3 per cent, contraction for
the rotor ducts, 2 or 3 per cent, for the stator ducts and perhaps 3 or 4 per cent,
for some slots in the pole face, and only in those cases when he knows that the
nature of the slots necessitates a careful calculation of the contraction ratio does
he refer to the figure at all.

The curves given in Fig. 36 refer to the case where open slots are used.
Where the slots are of the form commonly found in induction motors with over-
hanging lips, the values of Kg are higher particularly in those cases where the

W.M. E

66 DYNAMO-ELECTRIC MACHINERY

mouth of the slot is wide and the air-gap short. If the tooth has the general form
illustrated in Fig. 42, we may lake the values of tr from the dotted curve given in
Fig. 35. In this case it is the mouth of the slot that we must multiply by o-.

•Hft

length, of gap ~ y

FIO. 36. — Ciurn [or qolcUy BmllnB tbe coDtnictlon ntio K, lor open alota.

The contraction ratios with semi-closed slots of the shape depicted
in Fig. 42 for different values of s and g can be ohtained at once from
Fig. 37. It will he seen that the values do not differ bo widely from those
given for open slots as to make it worth while to consider intermediate shapes
of slot.

THE MAGNETIC CIRCUIT 67

An example of the uee of these curves is given on page 417.
The armature wiadings of nearly all modem alternating-current and
continuous-current machines are placed in slots in an iron core. In some cases the

wtS

width of slot g
lenffth ^ gap S

Fia. IT. — Correa for quJcklj' flDdlnj tbc contractton rtUo Kf t

field windings are so placed. The main reasons for placing windings in slots are :

(1) To give mechanical support and the protection of an iron-clad surface.

(2) To transfer to the iron teeth the forces which would otherwise come upon

the conductors.

(3) To reduce the reluctance of the magnetic circuit.

68 DYNAMO ELECTRIC MACHINERY

The circumstaDces which settle the size and shape of slots vary according to
the class of machine with which we are dealing. As each class of machine is con-
sidered in its place,* these circumstances will be considered at length.

In general, if we wish to make the ampere-wires per inch great, we would like
the slots to be large and roomy. On the other hand, to increase the magnetic loading
of the machine, we should like to make the teeth wide and the slots narrow. These
two courses being inconsistent with each other, we must make a compromise, and
considerable judgment is required to make the best use of the room at our disposal.

Choice of number of slots. In high- voltage machines (10,000 volts and
upwards) the slots will be made as few as possible and as large as possible, so that
the space occupied by the thick containing walls of insulation shall not take up
too great a proportion of the total space available for the winding. The drawback
to having too few and too large slots is that the cooling surface of the coils will be
small compared with the total cross-section of the coils. Moreover, the wave-form
of the machine may be prejudicially affected.

In low-voltage machines, where the insulation does not occupy so much room,
the tendency will be to increase the number of slots, so as to obtain a large cooling
siurface. In C.C. machines, the number and size of the slots are usually settled by
the commutating conditions (see page 479).

Depth of slots. The question naturally arises as. to what fixes the depth of
the slots. In c.C. and a.c. generators, it is usually necessary to preserve some
stated ratio between the ampere- wires on the armature and the ampere-turns on the
field-magnet. Very frequently the ampere-turns on the field-magnet are limited
by the copper and iron space available, so that a limit is fixed to the ampere-wires
per inch on the armature. It is therefore not necessary or desirable to make the
slots any deeper than is sufficient to accommodate the copper which will carry this
electric loading. Where the ampere-wires per inch of periphery are not limited
by considerations such as these, it is possible to increase the depth of the slots
until the leakage across the slots on full load begins to bear too great a ratio to
the working flux of the pole. This ratio of the leakage flux to the working flux is
the main consideration which determines the depth of the winding space in field-
magnets. For instance, in the case of revolving field-magnets for engine-driven
alternate-current generators, it is found that no advantage in output is to be
obtained by making the radial length of the poles more than 2^ times the width
of the pole, because the increased leakage between the poles neutralizes any
advantage that we can obtain from the increased winding space. This figure for
the ratio between the radial length of the pole and the width of the pole of course
differs in different circumstances (see page 300).

When we come to consider the field-magnets of A.C. turbo-driven generators of
the cylindrical type, we shall see that the limit in depth of the slots is sometimes
fixed by the amount of room which it is necessary to leave behind the slots. The
copper space on the rotor being thus limited, the ampere-turns on the armature
are likewise limited, and the depth of the armature slots will be made just great
enough to accommodate the requisite copper conductors. Otherwise, in the case of

* For slots of A.C. generators, see page 322 ; of induction motors, see pages 422 and 453 ; of
C.C. generators, see pages 480 and 490; also page 533.

THE MAGNETIC CIRCUIT . 69

an external annature, there is hardly any limit to the depth which mij;ht be chosen
for the annature slots. The greater depth would, of course, increase the armature
self-induction, but very often it would be an advantage to have this increased.

Id induction motors, the depth of the slots is an important factor in determin-
ing the leakage flusf, on which the performance of the motor greatly depends (see
page 422).

Very often the amount of electric loading on an armature is limited by the
cooling conditions on the end connectors. A deep slot will often necessitate an
arrangement of copper on the end connections which makes the cooling difficult.
This is a consideration which sometimes makes it desirable to use a shallow slot

Id direct current machines, the possible depth of the slot is sometimes deter-
mined by the commutating conditions, and there is very little doubt that where
commutating poles are used, the slote can be made much deeper than in machines
without commutating poles.

Owing to the above considerations, one will generally find in practice that
the depth of an armature slot is not greater than one-fifth of the pole pitch, and
in good regulating machines of conservative design, the depth is often not more
than one-tenth of the pole pitch.

Fonns of slots. Opsn slots. It is very convenient in many machines to form
the coils beforehand and put them into open slots. Where the coils are secured

by banding, the slots, may be made of the simple rectangular form shown in
Fig. 38. Where it is intended to secure the coils by means of a wedge of paper
or wood, the slot may have any of the forms shown in Figs. 39, 40 or 41. The
form in Fig. 39 has the advantage of requiring only a small and simple form of
wedge. The form in Fig. 40, however, gives a stivnger wedge for the space
occupied, and is for that reason ofton used in turbo-generators. Where the slot is
very wide and a still stronger wedge is required, the form shown in Fig. 41 is
useful. In this case, the coil is too wide to be inserted as a whole. It commonly
coDsists of two or four conductors, or it may be two or four coils, each of which is
inserted separately.

The magnetic leakage across an open slot will be smaller than the leakage
across a semi-closed elot, but^ as we have seen on page 65, the reluctance of the
air-gap is greater with an open slot, and the loss in the pole Face is greater,
particularly when the air-gap is short.

70 DYNAMO-ELECTRIC MACHINERY

Semi-dosed slots. For short air-gaps, and in cases where it is required to
reduce the ampere-turns to a minimum, semi-closed slots, sucfa as illustrated in
Fig. 42, are widely used. The addition of the lips makes the process of winding
much more difficult. Nevertheless, the great advantages to be obtained in some
induction motors and some alternate-current generators has brought the semi-
closed slot into ver; wide use.

As slots are placed around the periphery of an armature, as shown in Fig. 43,
their medial lines mm cannot be parallel, and on small armatures there is often
a very considerable angle between one slot and the next. The question, therefore,
arises whether it is better to make the sides of the slot parallel and the sides of
the teeth converging, or vice versa. In cases where the slots are open and
adapted to take coils of rectangular section, slots with parallel sides will be
employed. In this case the sides of the teeth will not be parallel. ^VIle^e the
armature is the revolving element, they will be narrower at the root and wider at
the periphery. Where the armature is the stationary element, they will be wider

FiQ. 43.— ShowtngboirpanUslslotalMd toUpeiteetli, Fio. 44.— ShowlnE pinllel taetb mi taper

eapeclallT wbne ths slota m tev In nnmber •Dd luge aa elati. uid the muuei ol atUUlnc the ifix In
comiMnd wltb circle on which they an placed. taper slota.

at the root and narrower at the periphery. From the magnetic point of view,
teeth narrow at the root are not as good as parallel teeth. The flux at the root
is generally as great as or greater than the flux at the periphery, and the ampere,
turns required on a taper tooth when highly saturated are very much greater than
for parallel teeth of the same mean cross-section carrying the same flux. This
will be seen from the example worked out on page 76. To make the best use,
therefore, of available space, one would prefer to use parallel teeth and slots with
converging sides, but the difficulties which this arrangement would ordinarily
entail have led to the more common use of parallel slots. In those cases, however,
where mush windings (i.e. windings consisting of a large number of small wires
placed haphazard in an insulated slot) are used, and where the exact form of the
coil is of little importance, considerable advantage can be obtained by employing
a parallel tooth and tapered slots (see Fig. 44). Again, in the case of bar-wound
armatures or field-magnets, where there are comparatively few bars per slot, it
may in some cases be an advantage to shape the bars so as to fit into tapered
slots. Such an arrangement is shown in one of the slots in Fig. 44.

The floxdensity in the teeth. When the teeth are highly saturated a con-
siderable portion of the flux finds its way down the slots and the ventilating

THE MAGNETIC CIRCUIT 71

ducts, ao we must consider the teeth, Blots and ducts as constituting magnetic
paths in parallel. For Bhortness of expression, we shall speak of the teeth as
" iron," and the slot space, duct space and space occupied by the insulation
between the iron sheets as "air." In Fig, 45 we are supposed to be looking
down on the top of slots. We can draw a rectangle ABCD around the space
occupied by one tooth and one slot between two yentilating ducts and as much
of a duct as lies within one tooth pitch. In Fig. 45 a small rectangle is
portioned off to represent the space occupied by the paper or other insulation.
The ratio of the whole area ABCD to the
area of the iron EFCG we shall denote by

A',= — i . On a machine of laree diameter,

iron *

where the sides of the teeth are nearly

parallel, ^i is almost a constant for any

section through the teeth.

On small armatures with teeth much

tapered, as in Fig. 43, it is very far from

constant, being perhaps 10 per cent, greater

for a section through the root of the teeth

than for a section through the tops of the

teeth. No very great error is introduced in

regarding K, aa % constant if its value is

calculated at a point distant by one-quarter ^°' *^,7f'^^''d'^ ^^4^^3*5? <^*-

of the length of the tooth from the narrowest

part of the tootb. Thus, the method of calculating K, for a revolving armature is

as follows: From the diameter da subtract 1^ times the length of the teeth I,, and

multiply by ff to get the mean circumference jr{da-l-5li). From this subtract

the number of slots multiplied by the width of the slots, and multiply the

difference by the net length of iron. This gives us the total section of iron in

all the teeth. Then multiply the quantity x(rfo - 1-5/j) by the gross length of the

armature core to get the total section of iron and air. The ratio K, is then

obtained by dividing the latter section by the former section.

ESANPLxll. AoontinuouB-ouirent armature U 120"diain. and ia 13" long. It hoa 324 Biota,
each i" wide and 2' deep. There are 4 ventilating duote, each |" wide. What is the section
of all the teeth and the value of K,, taking the solidity of the iron as 89 % T

Net length of iron (13 - l'e)0'8g= 10'2,

206 X 10-2 = 2110 aq. in. of aoUdiron,

368>! 13=4784 eq. in. of air and iron,

In the above example, anppoae that A B" = 300 megalines. Then the apparent flux-detisity
nn the teeth will be 300 x 10"t2100= 143,000 lines per sq. inch.

The stale of satonttion of the teeth will in this cose be so high that a considerable percentage
of the Sox will find ita way by the slots and dnote.

DYNAMO-ELECTRIC MACHINERY

The curves in Fig. 46 show the relation between the actual flux-density in the
iron and the apparent flux-density for different values of Kg- These curves are
easily plotted for any magnetic material as follows: Write down from the

■

■8

S
2
S

•*»

§ I

magnetization curve a list of the values of B and H. Draw on squared paper the
line for Kg=\ at 45** through the origin. If jr«= 1 there are no slots or vents, and
the armature is the same as if of solid iron ; the apparent flux density is equal to
the actual flux-density. Now, for JT, = 2 we have the air space in Fig. 45 equal
to the iron space, so that when the actual B = 22,000 and H = 800 the slot space

THE MAGNETIC CIRCUIT 73

carries 800 lines per sq. cm., which, when added to the 22,000, gives us 22,800
apparent flux-density. When plotting the curve for ir« = 2, therefore, it is only
necessary to add the values of H to the abscissae of the curve for Kg— I, Similarly,
the curve for Kg =^1-5 is obtained by adding half the values of H to the curve for

Example 12. An armature is 200 oms. in diam. and is 25 cms. long. It has 288 slots, each
1 *04 oms. wide and 4 cms. deep. There are three vents, each 1 cm. wide. What is the section
of all the teeth and the value of K$ ?

(200-6)t = 610

288x1-04=300

310 X (25 -3) =6850

6850 X 0*89 =6100 sq. oms. of iron,

610 X 25= 15,250 sq. cms. of air and iron,

16250
^•'■6100"^^-

Suppose now that the total flux of the frame A9B=140 megalines. Then the apparent
flux-density in the teeth is j^ ^ 10«- 6100 =23.000.

To find the actual flux-density from Fig. 46, follow up the perpendicular from 23,000 apparent
B to the 2*5 line, and we get actual B = 21,900.

Where the diameter of the armature is great as compared with the size of the teeth, so that
the sides of the teeth are nearly parallel, it is sufficient to calculate B in this way, and from the
magnetization curve find the ampere-turns per centimetre, which, when multiplied by the
length of the tooth, gives the ampere-turns on the teeth.

For instance, in the last example, the ampere-turns per centimetre for a flux-density of
21,900 in solid iron are 590 (see Fig. 47). The ampere-turns on the teeth are

4x590=2360.

Where very high densities are employed it is convenient to have curves which
gives us directly the ampere-turns per inch or per cm. for any apparent flux-density
and any given Kg, such curves are given in Fig. 47.

The method of calculation given in the last example would lead to inaccurate
results if applied to those cases in which the teeth are very much tapered and
fairly highly saturated, because a small increase in the flux-density at the root of
the tooth may call for a very great increase in the ampere-turns per centimetre.
Some designers take several sections of the teeth at different distances from the
root and calculate separately the ampere-turns necessary to drive the flux along
each portion of the tooth.

A method * which is less tedious than this, and at the same time more accurate,

is the one employing the curves in Fig. 49, which show how the values of I H £^B change

with the values of B for different values of Kg, The method is founded upon
the following theory :

Imagine a very long taper tooth (Fig. 48). Fix a datum mark DD, from which
to measure lengths along the tooth, such as l^ and l^. This datum mark may be

* See Hird, Jour, Institution of Electrical Engineers^ vol. 29, p. 933.

DYNAMO-ELECTRIC MACHINERY

somewhere at the wide end of the tooth, where the flux-density is very low ;
Z^ and lo are lengths measured from the datum line towards the narrow end.

200 4C0 eoo aoo 1000 aoo f4Co teot? moo 3000 zsoo 3400 moo 2900 3000 S200

1000 2000 SOaO ^000 5000 6000

Ampere tmyis per inch.

7O0O

aooo

Fio. 47. — Cmves giving relation between apparent flux-densities in teeth and the ampere-tums

per unit length for different values of Ku

For small changes in / we may take B as almost following a linear law. This
is the more true in practice where K^ increases at the root of the tooth. Thus we
can write approximately B = A;Z + constant, where A; = (Bg - Bj) -^ (/g - Z^). Bj and B.^
are the flux-densities at the distances \ and \ ^I'om the datum mark.

Fio. 48. — Showing convention upon which the rules for dealing with taper teeth are based.

THE MAGNETIC CIRCUIT

Thus, we have dS^kdl

Now the magnetomotive force required to drive the flux from A to C is

r* 1 f ^ ,

Hrf?=- HdB.
Jli «^Jbi

- &

1 1 § § i § § § i § § § I § § § § § § i § i

1 I I I I I I I t 1 I I I I I I 'I I 1 I 1 ' I'

I I » I I I I I

'H9NI OBd SNdrU dNV X NI'DS iiad SaNH ddVX

o
£

If, therefore, we have plotted curves of which the ordinates give the values
of I H dB, the abscissae representing B, the value of I HdB can be immediately

76 DYNAMO-ELECTRIC MACHINERY

obtained by subtracting the ordinate for Bj from the ordinate for B^. And

the ampere-turns in the tooth will be ^ — ^ I H dB, provided that we have

°2"°iJbi

employed suitable units in plotting the curves.

In Fig. 49 we have given the value of iHdB for the cases where the slot

and vent spaces form parallel paths with the teeth. Each value of Kg requires
a separate curve. It is sufficient to plot curves for the values of JT, given in
Fig. 49. The positions of intermediate curves can be judged very well by eye.
The manner in which these curves are employed to find the ampere-turns
expended upon the teeth is best seen from an example.

Example 13. The armature of a direct-current motor is 36 cms. in diam. and 25 cms.
long. It has 37 slots, each 1*1 cms. wide and 3 '5 cms. deep. There are 2 ventilating ducts,
each 1 cm. uide, and the iron laminations are 91 % solid iron. What number of ampere-turns
is required to drive the flux along the teeth and slots when the total il,B=26 megaliues
(see page 7)?

First find the cross-section of all the teeth on the tops.

36t = 1]3
37x1-1= 40-7

72-3 X (25 -2) = 1660x0-91 = 1610 sq. cm. of solid iron in

tops of the teeth.

Apparent flux-density — rgY^ = 17,220 =Bi,

113x25=2820 sq. cms. of air and iron,

2820
K, for tops of teeth =YFT^= 1-87.

Next take the roots of the teeth

(36-7)ir=91
37x11=40-7

50*3 X (26 -2) = 1160x0-91 = 1055 sq. cms. of solid iron in

roots of the tooth.

26 X 10*
Apparent flux-density -^^^^-=24,650=82,

91x25=2280,

2280
K$ for roots of teeth =77^^== 2 -16.

lUoo

Referring now to Fig. 49 and taking the curve for K, =2,

For B,= 24650 the ordinate =3 -2 xW.

For Bi= 17220 „ =015xl0«

7430 305xl0«

3-5
3-05 X 10* X =25^=1440 ampere- turns on the teeth.

The quantity /|2 - /j is of course the length of the tooth, in this case 3*5 cms.

It should be noted that in this book we consider the ampere-tunis on one
pole, not the ampere-turns per pair of poles as is sometimes done.

Air-gap and tooth saturation curve. The name '^ saturation curve'' is
sometimes given to a curve which shows the relation between the voltage
generated by a machine and the exciting current, or to a curve which shows

THE MAGNETIC CIRCUIT 77

the relation between the flux per pole and the ampere-turns on the pole.
Curves of this kind are given later (see pages 365 and 398). At this stage we
wish to consider another kind of saturation curve, namely, one showing the
relation between the flux-density in the air-gap and the ampere turns on the
^p and teeth. Such a curve is of the greatest ser^'ice in all investigations of
the flux distribution under a pole on no load and on full load.

It will be seen, in the first place, that for a certain armature punching, built
up with a certain solidity and with a certain number of ventilating ducts,
there will always be certain relation between the flux-density in the air-gap
^nd the saturation of the teeth near the region in the air-gap, where the flux-
density is under consideration, and this is independent of the number of poles or
the state of the load. Thus, a certain number of ampere-turns will always be
required for gap and teeth (taken together) for a certain flux-density in the gap.
In what is said here we are of course neglecting the irregularity in the flux-
density produced in the immediate vicinity of a tooth considered by itself, by the
presence of an open slot or any such very local disturbance. That disturbance is
Allowed for in the contraction ratio, but otherwise is neglected. By flux-density
in the gap we mean the average flux-density over the pitch of one tooth.

In plotting a gap and tooth saturation curve it is convenient to compare
all flux-densities to the density at a point in the air-gap mid-way between
■armature and field-magnet. If we start with the quantity Ag^ (see p. 7),
■and divide this by Ag^ the area of the active surface of the armature, this
surface should be taken to be the cylindrical surface lying mid-way between
the armature and the field-magnet. Thus, with a rotating armature

Ag^TF{ia-^g)y^la'

The apparent flux-density in the teeth at any distance from their roots is
obtained by dividing Ag^ by the total area of all the teeth at that distance
from the roots.

To plot our curve, then, we want in the first place ^^ as a standard of
reference for all other areas through which the flux has to pass.

Now take B = 10,000 (or Bjr=10 kapp lines per square inch if we prefer
those units), and calculate the ampere-turns on the gap, making allowance for
the contraction ratio as was done on page 65.

On a piece of squared paper lay out ampere-turns per pole as abscissae, and
flux-density in the gap as ordinates. As the ampere-turns on the gap are strictly
proportional to the flux-density in the gap, a straight line joining the point giving
the ampere-turns on the gap for B= 10,000 with the origin will give the ampere-
turns on the gap for any flux-density. It is known as the air-gap line. Now take
several values of B in the gap, say, 8000, 9000, 10,000, and 11,000. For each of
these values divide Ag^ by the section of all the teeth, and calculate the ampere-
turns required for the teeth for each value. Lay off on the paper those additional
ampere-turns as additions to the abscissa for the ampere-turns on the gap for each
value of B in the gap. This gives us the curve we want.

Example 14. Take the data giveu in Example 11, page 71. Assume that the radial length
of the air-gap is I", and plot the gap and tooth saturation curve.

DYNAMO-ELECTRIC MACHINERY

To get the ampere -turns on the gap we must first take the contraction ratio.

Now, 5 = 0-5 and ^ = 0*25, - = 2. Pitch of slots is M6=p„ so that " = 0-43.

From Fig. 36 the contraction ratio due to slots is 1135.

There are 4 vents, each f wide, total lb" The armature is 13* long, so that

— for the ducts is ^r77 = 0115, and - for the ducts =^r-^^ =1'5. Kg for the
Ps 13 ' ^ 0-25 ^

ducts = 1 '03. The total contraction ratio is therefore 1*135 x 1-03 = 117.

For B" = 60,000 lines per sq. in.

Ampere-turns on the gap = 0*313 xB" xg" x Kg

= 0-313 X 60,000 X 0-25 x M7 = 5500.

70000

60000

50000

1^ 40000

■•A

«^

30000

20000

10000

«.•.----

5500

on gap

4
/

SOonteet

^.^-^--^

/ yT

fOOO

7000

8000

2000 3000 4000 5000 SOOO

Ampere Turns on Gap and Teeth

Fig. 50. — niuBtrating the method of constructing an alr-gap-and-tooth-eatoration carve.

Mark on squared paper the point B" =: 60,000, Kg = 5500, and join the point to
the origin. This gives us the air-gap Una

As the teeth are short as compared with the diameter of the armature, it is
sufficiently accurate to take the area of the teeth as we did on page 71.

At = {{d - VUt)Tr - (No. of slots X width)} x net length of iron.

In this case At = 2370 sq. in. of laminations, or 2110 sq. in. of solid iron.

Now ^^ = (120 + 0-25)^ X 13 = 4900 sq. in.,

^' = 4900 = ^^^-

THE MAGNETIC CIRCUIT

To get the apparent flux-density in the teeth for any flux-density in the gap
we merely divide by Kty and to get the actual flux-density we can refer to Fig. 46.
As we know Kg (in this case 2*28, see page 71), we can refer at once to Fig. 47 and
read off the ampere-turns per inch required for the teeth.

Make four columns :

B" in gap.

B" app. in teeth

Aiapere-tunia

Ampcre-tumB

=B^'~0'48.

p«r inch (Pig 47).

iu teeth.

45,000

105,000

120

50,000

116,000

180

360

55,000

128,000

400

800

60,000

139,000

850

1750

65,000

151,000

1900

3800

It is hardly necessary to remind a technical student that to get column 2 from
column 1 it is only necessary to put 4*3 on the C scale of a slide rule opposite 1,
and then read off column 2 from scale D,

The higher values of the ampere-turns per inch can be taken from Fig. 47.
The lower values can be more accurately taken from Fig. 21. The additional
ampere-turns required for the teeth are plotted as in Fig. 50. Other examples of
curves of this kind will be found in Figs. 301 and 373.

The considerations to be kept in view in designing the tips of teeth on
annatnre cores. These are as follows:

(1) The object of the tip is to make the head of the tooth as wide as possible
without increasing unduly the inductance of the conductors in the slot.

(2) Sufiicient iron must be provided at the root of the tip to carry the flux
passing through the tip, and to give it mechanical strength.

(3) The permeance of the magnetic path encircling the slot (that is the path
for leakage lines) is to be kept as low as possible.

(4) The slot should be of such a shape as to only require a simple die to punch
it, and the comers should be such that they will punch well without requiring the
frequent repairing of the die.

(5) The mouth of the slot must sometimes be not less than a certain minimum,
as when it is intended for mush coil winding.

(6) The method of drawing the slot and making the punch for it should be
capable of being easily standardized.

The following names of parts and symbols will be used (Fig. 51 illustrates
the parts):

A = height of slot. m = mouth of slot.

Ac = height of conductors. p = \ip of mouth.

6 = width of slot. ^ = air-gap.
r = root of tip. a = angle of slope of tip.

As to the general shape of the tip, it is only in special cases that anything
is to be gained by the use of a large radius at the comer of the root, as in

DYNAMO-ELECTRIC MACHINERY

Fig. 52. For a standard slot, of normal size {b from 0*25 to 1 inch, k from 0*5
to 2 inches), intended to take various numbers of conductors, and various amounts
of insulation, the shape of slot (Fig. 51) is as good as any other. In drawing
it, the comers may be shown sharp, and the die maker will put on a very
small radius to get good results in punching. He can make a cheaper tool
than if he has to work to special radii which change with every slot.

In general, the best value for the angle a is about 27 degrees, that is,
tan~i0*5. This angle gives sufficient iron in the root of the tip both for the
working flux and the leakage flux, when the flux-density in the gap is as high
as 60,000 lines per sq. inch. The angle might in some cases be reduced where
it is very desirable to save space, but in general the same effect can be obtained
by drawing the sloping line a little lower down, while still keeping it at the
same slope as shown in Fig. 53. For slots of the most ordinary size the apex
of the angle a may lie on the centre line of the slot as shown in Fig. 53.

^ ^

Fig. 52.

FZG. 63.

The dimension m may be fixed by the necessity of putting wires of a certain
size through the mouth of the slot. In any case it should not be made too
small (say not less than 0*05 inch), on account of the necessity of giving sufficient
strength to the metal of the punch. Subject to these considerations m will be
made as small as possible, so that the face of the tooth may be as large as
possible, due regard being had to the effect of the shape of the tip on the
permeance of the path for magnetic lines immediately encircling the slot.

It is useful to have a diagram like that given in Fig. 54, which gives at a
glance the values of permeance of the magnetic path across the mouth of the
slot for different shapes of tips. This diagram is used in the following way :

Suppose that the mouth of the slot, ?7i, is to be 0*35 of the width, b. At the
point 0'35 on the horizontal scale erect a perpendicular as shown in the figure.
If i.t has been decided provisionally that the apex of the angle a shall lie on
the centre line of the slot, this perpendicular is drawn to meet the line OA^
and the shaded area gives us a picture of the tip. As the perpendicular is
always one half of the abscissa, we can judge at once whether the dimension p
is or is not too small to punch well. Now carry up the perpendicular (as
shown by the dotted line) until it cuts the curve A\ and the ordinate gives us

THE MAGNETIC CIRCUIT 81

the value of the permeance of the path acroes the mouth of the slot for one
centimetre length of iron, independently of the size of the slot. For instance,
in the case taken in the figure there would be 0*98 c.G.s. lines across the mouth

of the slot, for every cm. of length of slot, for every j- amp. carried by the

slot. In order to get the total permeance of the slot, this value must be added
to the permeance of the path between the parallel sides. If for any reason it is
desirable to lower the sloping line, as in Fig. 53, then a, line such as £ in
Fig. 64 will be the boundary line of the slot, and the curve ff gives the per-
meance of the path across the mouth of the slot. If, as in turbo-machines,

0-1 O-I B-l ft O-t W 0-7 u-g 0-S II)

FW. M.— CorvM loi quickly ealoilalliig ttaa penwuioe of Ui* path ktou the montl) <il ■ ilot.

it is necessary to make the tip thicker at the root, the sloping hne may be taken
ID a position such as DC, then the permeance is given by the curve C. For
any intermediate size or shape of slot it is easy by eye to interpolate the point
on an imaginary curve, say between A and B, which gives the value of the
permeance of the path across the mouth of the slot.*

For instance, to calculate the leakage flux per centimetre of axial length of iron
for the slot shown in Fig, 51 when 200 amperes are flowing in the conductor, we
proceed as follows : The leakage across the body of the slot for one ampere in the
conductor is i^ \ h 9.7

*' ' * =-419x?r! = 0-87.

10 3

1-3

82 DYNAMO-ELECTRIC MACHINERY

The leakage across the mouth of the slot when

I « '^1 = -35 is 0-89 (from Fig. 54).

The total leakage for one centimetre axial length of iron for 200 amperes is

(0-87 + 0-89) 200 = 375 c.G.s. lines.

When we are dealing with an alternating current we must remember to take
its maximum value if we want the maximum value of the leakage.

For examples of the calculation of the leakage across slots, see pages 422
and 463.

Flnz-densities in the teeth. In slow-speed continuous-current machines, where
the frequency is low (15 to 25 cycles) very high flux-densities in the teeth can be
employed. There is an advantage in employing high flux-densities in such cases,
as the commutation is improved thereby, and it will be seen from Fig. 29 that at
low frequencies very high saturations can be employed without danger of over-
heating. Flux-densities as high as 21,000 c.G.s. lines per sq. cm. can be employed
with advantage on such machines, and, allowing for the amount of flux that finds
its way through the slot space and ventilating ducts, the apparent flux-density
may be as high as 28,000 c.G.s. lines per sq. cm. (see Figs. 45, 46 and 47).

Similarly, in 25-cycle A.c. generators and induction motors very high flux-
densities are often employed in the teeth. A density of 22,000 is not uncommon
in such cases, but the cooling conditions of each case must be studied to see if such
densities are permissible (see page 324). In the case of induction motors the density
is often limited by the prescribed limit to the magnetizing current. This is especially
so on machines having a small pole pitch (see pages 419 and 446).

In 50-cycle machines it is generally necessary to reduce the flux-density so
that the losses on the teeth may not be so excessive as to interfere with the coolingr
of the coils. A density of from 18,000 to 20,000 may be taken as fairly high for
50-cycle machines. Each case must be considered with regard to the cooling
conditions (see page 470) and effect on the efficiency.

The iron behind the slots. For the types of machines ordinarily manufactored
it will be found that it is not worth while to calculate accurately the number of
ampere-turns required to drive the flux through the armature core (or the
iron behind the slots, as it is sometimes called). The reason is that in mosl^
cases these ampere-turns are small compared with the total ampere-turns oix
the pole, so that an inaccuracy of 50 per cent, will hardly afiect the total. As
the flux distributes itself in some such manner as indicated * in Fig. 58, it would

be necessary to find |J7(S all along la (see Fig. 31) in order to find the ampere-^

turns correctly. As this is too much trouble in practical calculation, one adopte
the following rule, which, though far from giving an accurate result, is good enough
when one considers the uncertainties that enter into more important parts of the
calctdation of a machine. Find the maximum flux-densitv in the iron behind the
slots by dividing the working flux per pole by twice the cross-section of the iron

♦See Dr. W. M. Thornton, "The Distribution of Magnetic Induction and Hysteresis Loss.
in Armatures,'' Jour. Inst. Elec. Engrs., vol. 37, page 126.

THE MAGNETIC CIRCUIT 83

behind the slots. The ampere-turns per pole required to drive the flux thmugb
the core will in general be found to be rather leas than the ampere-turns required
to create this flux-density in an iron path, whose length is equal to one-thiid of
the pole pitch. So, if we find the number of ampere-tuma per centimetre required
for the mazimom flux-density in the core, and multiply by one-third of the pole
pitch in centimetres, this gives ua a aafe figure for ampere-tums.

Fia. SG.— DUtributlon ol flux la the iTon behind tb« teeth.

Example 15. In the 1600 h.p. motor worked out hi Chapter XVII., the flux per pole is
5-6x]0'co.s. lines. The section of iron behind the sloU ia 332 sq. oma. The flux^dengity
ia therefore 5-6x 10'-Jfl64 = 8450=B.

This will require about 3 tmpere-tuma per cm., and tta the pole pitch ia 31*2 cms., the
ampere-tnniH per pole required for the core are about 32. As the total ampere-turns per pole
are over 1400, it will be seen that it would he useleas to mabe a more careful estimate.

In small armatures less than 36 inches in diKmeter, the punchinga are usually
made in one piece, so that the iron behind the slots forms an unbroken magnetic
path of very low reluctance. These unbroken cores will carry a greater flux for
a given iron loss than cores built up of interleaved segments, baving many breaks
in the circumference. The breaks in the continuity of the iron bring about losses,
not so much at the breaks themselves, as in the surrounding parts of the iron core
and &ame, owing to the reluctance of the break. If the mean flux-density in the
core is as high as 12,000 or 13,000 lines per sq. cm., we will find that at the break-
joints part of Uie flux only keeps to the iron path, and some crosses the small air-
gap between the abutting ends of the broken punchings. The amount of the flnx
which crosses this air-gap is easily calculated ftom Fig. 16. Let the mean flux-
density in the core be 13,000. Then, assuming that one-half the punchings bridge
the break-joint, the apparent flux-density in the iron will be 26,000. From the
curve K,=2m Fig. 46 we find that the actual flux-density ia only 23,700, so that
the density in the air-gap will be 2300. If, now, the distance between the abutting
ends of the punchings is 0-02 inch, or, say, 005 cm., the ampere-tnms required for
the break-joint will be 2300x005x0-795 = 91.

Now, if, as depict«d in Fig. 28, the small air-gap between the breaks is not
uniform for the whole width of the core, there will be a tendency for the flux to

84 DYNAMO-ELECTRIC MACHINERY

crowd to the side of the machine where the air-gap ia smallest, and it will be seen
that 91 ampere-turns, or even half of it, is sufficient to produce a considerable
difference in the distribution of the flux.

Even when this small air-gap is uniform, there is a tendency for some of the
flux to be driven out into the iron frame or end plates, but very little loss can
occur from eddy currents due to this cause, unless the gap is too big or the flux
density in the core excessive. It is important in machines of low frequency,
where the core densities are made very high (14,000 to 15,000) to see that the
joints are well made. In cases where the frame is split and all the punchings are
cut through, the length of air-gap between the punchings is of vital importance.
The surface of the joint should be carefully finished off, so that the gap is reduced
to at most a few thousandths of an inch, otherwise the flux will be driven into the
frame and considerable heating occur aroimd the joint. In four-pole machines
a complete break in the pimchings on a horizontal or vertical diameter is almost
sure to produce an eddy current in the shaft. Eddy currents in the shaft can also
be caused by dissymmetries in the break-joints of an armature, even when there is
no complete break in the punchings.

For a frequency of 50 or higher the flux-density in the iron behind the slots
is generally limited by considerations of iron loss. For low frequencies (15 cycles
or lower) the reluctance of the path is the more important consideration. We
will take first the higher frequency cases. The amount of iron loss permissible
in a core depends upon the facility with which the heat can be dissipated. If the
ventilating ducts are very near together, and there is a good draught, one can work
the core at a higher density than when the ducts are further apart and the draught
not so good. In 50-cycle machines with the natural ventilation obtained from an
ordinary speed, and having f inch ventilating ducts spaced with a pitch of 25 inches,
one can work safely at a core density of 11,000 lines per sq. cm. This gives, according
to Fig. 29, a loss of about 1 watt per cubic inch, and in a core 4 inches in depth
the cooUng surface comes out at about 1*2 sq. in. per watt (for core loss only),
counting both sides of the ventilating duct and the surface at the back of the iron.
When the frequency is higher, it is usual to increase the number of ventilating
ducts so as to be able to work at a higher loss per cubic inch ; and at the same time
the flux-density is diminished so as to keep the loss per cubic inch within reasonable
bounds. At 100 cycles, for instance, one might, ¥dth ordinary cooling conditioDB
and ordinary iron, work the core at 8^8000, giving a loss of, say, 1*5 watts per
cubic inch. The ventilating ducts might then be spaced with a pitch of 1*5 inches.
The depth of iron in high-frequency machines is for the same speed smaller than
in low-frequency machines, so that the cooling conditions are better. In turbo-
generators with forced draught, the permissible density in the core requires special
study (see page 391). In low-frequency machines it is not advisable to increase
the flux-density much above 16 kapp lines or B» 15,000, because at about that
point the ampere-turns per inch increase very quickly.

The flux pet pole is calculated by dividing the quantity A,Bhj the number of poles
and multiplying the firm constant K/. As the flux from the pole divides into two,
one-half going to the pole on the right and the other to the pole on the left, we must
divide the total flux by twice the area of the core to obtain the mean flux-density.

THE MAGNETIC CIRCUIT 86

Example 16. A direct-current generator has eight poles, with a flux form constant of 07.
The diameter of the armature is 92 cms. , and the net length of iron 28 cms. There are 06 con-
ductors in series, and at a speed of 375 B.P.M. it generates 255 volts. What must be the depth
of iron below slots in order that the density in the core shall not exceed 12,000 lines per sq. cm. ?

From formula (1), page 24, we have

255=0-7 X ^ X 96 X ^^B X lO-^.

^,B=60'7 megalines.

i?i 1 60-7 X 0-7 .00 T

I lux per pole ="5 — ^ — =o'32 megalmes.

o poxes

5320000 ^, ,. c

s — i75ww^=221 sq. cms. = cross-section of core.

2x 12000 ^

As the net length is 28 cms. the depth below is ^^=7*9 cms.

Example 17. A certain 4-pole a.c. turbo-generator frame has a punching whose depth
behind slots is 8|", and the net length of iron 40*. How many conductors must we have for a
3-pha8e star- wound generator running at 1500 b.p.m., in order that the flux-density in the core
shall not exceed 1 1 kapp lines per sq. in. ?

8f X 40=350 sq. in. of core,

350 X 2 X 1 1 = 7700 kapp lines per pole.

Take the form constant, A/, at 0*64 and the volt constant Ke at 0'4.

7700
AgBK=^fr:^x 4 poles =48,000 kapp lines,

6600 X 10«= 0-4 X 1500 X J?a X 48,000,

^a=229.

As 229 conductors would not be suitable for a S-phase generator, we might, if there were
48 slots, make 240. The density in the iron behind the slots would be 10*5 kapp lines per
sq. in.

Ampere-tiiniB on the yoke. In machines like continuouB-cunent generators,
having external field magnets, the length of path through the yoke is often quite
considerable, and the number of ampere-turns required for this part of the magnetic
circuit should be calculated with some accuracy.

The flux carried by the yoke includes the leakage flux as well as the working
flux, so that before a calculation of the ampere-turns can be made it is necessary
to calculate the amount of leakage. The graphical method of calculating the leakage
given on page 326 will be found to be very short and sufficiently accurate for
practical purposes. It has the advantage over the method employing formulae,
in the fact that it can be so easily adapted to varying shapes of pole. Moreover,
the designer, having a picture of the flux distribution before him, can more easily
check the result, and he can see what feature in the arrangement of the pole is
mainly responsible for the leakage.

Having determined the amount of leakage flux, this is added to the working
flux and the whole divided by twice the area of the yoke to get the flux-density.

Example 18. On a certain continuous-current generator the working flux amounts to
10*5 X 10* lines per pole and the leakage at full load amounts to 2*1 x 10^ lines. If the arrange-
ment of the yoke is as shown in Fig. 432, the area being 650 sq. cms., find the number of
ampere-turns required for the cast-steel yoke.

86 DYNAMO-ELECTRIC MACHINERY

Total flux = 12*6 X 10" lines. Divide by 2 and by 650, and we get 9700 C.G.S. lines per sq.
om. Referring now to Fig. 22, we find that this requires about 9 ampere-turns per cm., and
the effective length of yoke being 33 cms., we find the ampere-turns on the yoke to be about 300.

In the machine to which the above example refers, the cross-section of steel
has been fixed more by regard to the stiffness of the yoke than by magnetic con-
siderations, and the flux-density is much lower than one would find in generators
of smaller diameter. In a cast-steel yoke reasonably free from blow-holes one may
economically employ a flux-density as high as 12,500, and this usually requires
about 15 ampere-turns per cm. It is not good practice to carry up the saturation
much higher than this, because cast-steel is liable to have blow-holes in it, which
may cause unequal pole strengths in a multipolar frame, or may call for an excess
magnetizing current if the saturation is carried too far.

The following articles dealing with dynamo steel and iron losses are of importance :

" Hysteresis Loss in Induction Motors near the speed of Synchronism,** H. Zipp, EUhhrotech,
u. Maachinenbau, 26, p. 443, 1908.

" Iron Losses Induction in Motors due to Flux Pulsations," Bragstad and Frftnkel, EUktroL
ZeU„ 29, pp. 1074, 1102; 1908.

"Experimental Determination of the Hysteretic Constant,** N. Stahl, Elec WcM^ 52,
p. 1122, 1908.

" Best Thickness for Iron Sheets in Electrical Work,** Lopp6, Ind, Elect., 18, g. 413, 1909.

" Dependence of Magnetic Hysteresis upon Wave-form,*' M. G. Lloyd, Bureau of SUmda/rds,
Bull. 5, p. 381, 1909.

" Testing Iron by the Ballistic Electrodynamometer,** Rice and M'Collum, Pkys. Rev,, 29,
p. 132, 1909.

" Magnetic Testing of Iron with Alternating Currents,*' Campbell, InH, Eke, Eng,, Joum. 43,
p. 553, 1909.

" Calculation of Iron Losses in Dynamo-Electric Machinery,** I. E. Hanssen, Amer. I,E,E.,
Proc. 28, p. 679, 1909.

'* Hysteresis and Eddy-Current Losses in d.g. Machines,** Steels, Assoc, Ing. EL Libgt, BulL
9, p. 341, 1909.

"Comparison of Iron Losses during Alternating, Rotating, and Static Magnetisation,"
Czepek, EleHrot. u. Maschinenbau, 28, pp. 325, 351 ; 1910.

Iron Losses in a Rotating Field,** Hermann, Elektrot, Zeit., 31, p. 303, 1910.

Iron Losses in a Rotating Field,** Hiecke, EleHrot. u. Mtuchinenbau, 28, p. 683, 1910.

Magnetism of ' Stalloy,* *' H. R. Hamley and A. L. Rossiter, Roy. 8oc. Victoria, Proc. 23»
2, p. 325, 1911.

" Rotor Hysteresis in Polyphase Induction Motors,** D. Robertson, Electrician, 68, p. 12,
1911.

" Commercial Testing of Iron for Hysteresis Loss,** L. T. Robinson, Amer, I.E.E,, Proc. 30,
p. 825, 1911.

"Source of Extra Iron Losses in Rotating Smooth Ring Armatures,** J. Wild, ZeOsckr,
Vereines Deuisch. Ing,, 56, p. 1441, 1912.

" Variation of Magnetic Properties of Iron and Steel with Temperature,** Le Orensot, EleHrot,
u. MOfSchinenbau, 30, p. 986.

Dynamo Sheet Steel,** Metall. and Chem. Engin,, 10, p. 553, 1912.

Magnetic Properties of Dynamo Iron,*' De Nolly and Veyiet, Elektrot, Zeit,, 30, p. 985, 1912.

" Electrolytic Iron for Electrical Machinery,** Breslauer, Elektrot. Zeit., 34, pp. 671 and 705,
1913.

" Magnetic Investigation of Sheet-Iron,** Zickler, Elek. u. Maschinenbau, 31, pp. 737 and 759,
1913.

" Relation between Magnetic Properties of Steel and Temperature,'* Rev. de MitaUurgie, 10,
p. 146, 1913.

Mechanical and Magnetic Testing of Steel,** R. P. Devries, Rev. de MitaXl., 10, p. 141, 1913.

((

CHAPTER VI.

THE ELECTRIC CIRCUITS.

Annature windings. In laying out the armature winding of any dynamo-
electric machine, the logical procedure will be as follows:

(1) Lay out the condnctor diagram, i.e. the number of conductors, number of
phases, the position of the conductors relatively to the poles, and the direction in
which the currents will pass in the conductors at a particular instant. In this
scheme we are only concerned with these matters, and not at all with the end
connections.

(2) Make a connector diagram showing how the ends of the conductors are to
be electrically connected in order to carry out the scheme, thus obtaining our
winding diagram.

(3) Consider the mechanical design of the end connectors to see that they
clear one another with sufficient spaces between, and are mechanically strong
enough.

(4) Consider the material of the conductors.

(5) Their size and shape of cross-section.

(6) The effect of eddy cnrrents.

(7) Calculate the resistance and the weight.

(8) Consider the heating and cooling.

1 AND 2. THE CONDUCTOR DIAGRAM AND WINDING DIAGRAM.

Single-phase windings. Suppose that we are laying out the winding of a
single-phase generator : Lay out — on squared paper for convenience — fine dotted
lines to represent the pitch of the poles, as shown in Fig. 100. The lines of the
squared paper can conveniently be taken to represent the slot pitch. Draw with
thicker lines the conductors in their proposed relation to the poles, and show, by
arrow heads, the direction in which the current will pass at some particular
instant. In this scheme we settle how many slots shall be wound out of the total
possible number of slots in the pole pitch. For example, it is common in a
single-phase machine to wind two-thirds of the slots in the pole pitch, though, for
reasons to be dealt with when we come more particularly to consider the design of
such machines, another fraction may be taken.

DYNAMO-ELECTRIC MACHINERY

In Fig. 100 the pitch of the slot is one-ninth of the pole pitch, and we have six
wound slots per pole. The conductor diagram is therefore completely given in
Fig. 100. It is well to settle this simple matter first, before we proceed to the

>r >

' >

' >^

A J\ Ji A JK >

>r \f \f yr \f \f

Fig. 100. — Conductor diagram for diigle-phaBe winding.

winding diagram, because whatever end connections we may employ to carry out
the scheme, they will have no effect upon the amount of the e.m.f. generated, or
its wave-form.

We may, for instance, connect any one of the conductors passing under a
north pole to any one of the conductors passing under a south pole, and the effect

\f \^ \f \f \r \f

Ji A J\ A >^ A

V V V V \f \r

Fia. 101. — Concentric hemitropic connections for Bingle-phase winding.

will be the same, so far as it can be measured at the terminals of the machine.
There are, however, certain classes of end connectors which have been found in
practice to be the most satisfactory. These will be considered here.

End connections may be broadly classified into concentric connections and

THE ELECTRIC CIRCUITS

8»

lattice connections. The latter are sometimes spoken of as "overlapping" con-
nections. Figs. 101 and 102 show concentric connections; Figs. 103, 104, 105
and 106 show lattice connections.

It will be seen that these terms " concentric " and " lattice " refer to the type
of connections as shown in the "connector diagram." There are a great number
of different ways of carrying out mechanically each type of connection (see p. 115).

In Fig. 101 all the conductors lying under one pole are connected by means of
a broad band of connectors to all the conductors under another pole. This style
of winding has been called hemitropic*

In Fig. 102 the conductors lying under one pole are divided approximately
into two parts ; half of them are connected to conductors lying under one pole to-
the right, and half of them to conductors under the pole to the left. This type of

V V \f \r \f >f

Fio. 102. — Concentric connections for single-phase winding.

winding has the advantage in requiring less copper than the hemitropic winding,.
as the average length of the end connectors is shorter. It also has the additional
advantage in the fact that the armature reaction does not at any instant create a
difference of magnetic potential between the iron behind the armature slots and
the iron behind the poles, whereas with the winding depicted in Fig. 101 there is,
at the instant when the armature current is at its maximum, a very considerable
difference of magnetic potential between the armature frame and the field-magnet,,
which may cause serious eddy currents in the frame or in the shaft. In a three-
phase machine this effect is neutralized, because the total current at any instant
equals zero.

Where there are a great number of conductors per slot, these conductors will,,
in general, be grouped in a coil, their end connections being more or less parallel,
and they may therefore be considered as forming concentric connections, as
between themselves. The coils may then be assembled as concentric coils with the
connections between coils made either as in Fig. 101 or Fig. 102, or the coils may
be arranged as a lattice work in a manner somewhat similar to Fig. 103.

*See Polyphase Electric Currents^ by Prof. S. P. Tliompson, 1900, p. 85.

-90

DYNAMO-ELECTRIC MACHINERY

Fig. 103 shows a bar winding with lattice connectors, having a throw of a pole
pitch at one end and a pole pitch minus one slot at the other. Observe that in
this figure the winding is hemitropic, and its magnetic effect will be the same
JLS for the coil shown in Fig. 101. Such a winding should not be employed in
single-phase armatures.

Fio. lOS. — Lattice oonnectlons, the throw being a fuil-pole pitch at one end.

In making a diagram of lattice connectors, it is convenient to leave out most of
the connectors, as shown to the right of the figure. The diagram, besides being
easier to draw, is easier to follow, particularly when several phases are super-
imposed.

In Fig. 104 is shown a winding of lattice connectors, which in effect are the
same as Fig. 102. Here the mean length of connector is less than in Fig. 103, and
the magnetic action of the armature is symmetrical.

Fig. 104. — Lattice connections with short throw.

All the above connections result in what is sometimes termed a "lap"
winding as distinguished from a "wave" winding.

In a wave winding such as shown in Fig. 105, we pass under a north pole,
then south, and then the next north, instead of returning under a same north as
in the lap winding. Wave windings are very convenient to employ in a bar-

THE ELECTRIC CIRCUITS

wound machine, because by their use we do away with specially-shaped connectors
between one coil and another. It should be remembered here, that with a wave
winding the average length of end connector is greater than in the type of winding
fihown in Fig. 104. In wave windings it is usual to employ two conductors per

FIG. 105. — Lattice oonnections forming a wave winding. Six independent circuits doeed

on themselyee.

slot, as this arrangement makes it possible to have a symmetrical arrangement of
conductors on the two sides of the machine.

A few years ago it was usual with the simple (two-bar per slot) wave windings
to have an odd number of slots, so that after we had progressed around the
machine with a number of steps equal to the number of poles, we arrived at a slot

Fio. 106. — Lattice connections forming a wave winding with the throw of elz connectors
altered so as to put the six circuits of Fig. 105 in series and bring out two ends.

either one short or one ahead of the slot from which we started. Then we
stepped round the machine again, coming in either one short or one ahead, and so
on until all the slots were filled. This method had the advantage of calling for at
most two different lengths of end connectors, and it also had the advantage of
changing the phase of the slot very slightly at each throw. Such an arrangement
of slots is, however, not always convenient when a standard line of machines must

DYNAMO-ELECTRIC MACHINERY

be laid out. For a standard line, designed to be wound for many different
voltages, it is more convenient to have a whole number of slots per pole. In this
case the wave winding is just as possible as before, the only difference being that
after we have stepped around the machine with a number of throws equal to the
number of poles, we make one throw rather shorter or rather longer than before,
and come into a slot either one short or one ahead of the one we started from.
The number of special connectors required for this method is usually only small,
and the difference in pitch is so slight that it is hardly apparent after the machine

>r \f \f \f

>< J< J\ J<

Fio. 107. — Conductor diAgram for two-phase winding (fal

\f \f \f \f

pitch).

has been wound. The easiest way to lay out such a winding as this is to first of
all lay out all the connectors as if the throw were constant. We then obtain a
number of circuits closed on themselves, as shown in Fig. 105. For convenience
in tracing out the winding, we have affixed the numbers 1, 1 ; 2, 2 ; 3, 3,
etc., to distinguish each closed circuit where it leaves the diagram on the
right and where it begins again on the left. We then choose some convenient
part of the winding where we wish to put the terminals, and we shorten — or
lengthen — some of the connectors, as shown in Fig. 106, in a manner which puts

\f \f I \f

J< JK

>^ I

>r \f

Fio. 108. — Conductor diagram for two-phase winding (short chorded).

all the conductors in series with one another. It will be seen that by the shorten-
ing of the throw in the centre of the diagram, circuit No. 1 is put in series with
circuit No. 2, and so on.

All these figures (101 to 106) represent different imndvng diagrams for carrying
out the conductor diagram in Fig. 100.

Two-phase windinc^. In a two-phase winding, as before, first lay out the
condiictor diagram. Most commonly this will consist of a simple arrangement, such
as is shown in Fig. 107. This would be a full pitch two-phase winding. If the
winding were chorded, the scheme might appear as in Fig. 108. Fig. 109 shows
the arrangement of end connectors for this where there are two conductors per

THE ELECTRIC CIRCUITS

«lot. Other schemes of chording might be employed (see Fig. 120). As stated
before, these conductor diagrams do not concern themselves with the end con-
nections, though of course the scheme adopted will affect the length of the end
•connectors.

'<:)V'<>V

inro][o][T|T|"o|oj«|fF o

Fio. 109. — Showing two-phase wincUng alter the schame of Fig. 108 with two ban per dot

and connectors having a short throw.

Taking the simple conductor diagram given in Fig. 107, we may connect the
•conductors of phase A, just as if it were a single-phase machine, by any of the
methods illustrated in Figs. 102, 103, 104 or 106. Similarly, we may connect any

r r

f y

I ! I i

r • [ 1 1

< i

^ >

^ ' 1 I I

f y

r \

f >

r I « [

Fig. 110. — Conductor diagram for three-phase winding.

of the conductors in phase B in the same way, but we must remember that the
•connectors of one phase must keep clear of those of the other.

m m •

V ^a

^ ■

r. ....^

...

: fy

• ^ «

—

f \

/^ '

^ i

f \

r >

r y

r 1

I 1 [ J

i i

\ d

f \ \ \

\ \ \ \

; )

' >

r y

f 1

-■

«" '

1 : V...

'"y

...

—

•

L-

<s J

-J

FIO. 111. — Concentric connections on three-pliase winding in three tiers,

94 DYNAMO-ELECTRIC MACHINERY

In two-pbase machiDea a very commoD method iB to arrange the connectors in
two tiera, as illustrated in Fig. 113(o). The coils of phase A may have straight
ends as shown, and coils of phase B may be bent up so as to clear the projecting
bars of A. These are commonly spoken of as "bent ends." If the proportions
are approximately as shown in Fig. 113(a), the resistance and self-induction of
the two phases will be very nearly alike.

RearSnd-

Fia. 112.— Thne-plutM Tladbia oonsLBtlns ol concentric colli ■mnsad with Uu BDda Id two
Uen, inar tbs mknoer llliatntod In tig. 114.

Where lattice connections, such as in Figs. 104 and 105, are employed, the end
connectors of one phase lie contiguous to those in another, so that with these type»
of windings it is necessary to insulate the end connectors to withstand the full
pressure between phases.

mfid

Fio. lis. — Showing TutoDi mettaodi of untnglna ends of ooUb In ■

IT winding.

Three-phase windings. A conductor diagram for a simple three-phase winding
is shown in Fig. 110. The most straightforward way of making the end connectors
for this is to arrange them in three tiers, as shown in Fig. 111. These three tiers
may be arranged in the methods shown in Figs. 142 and 348. Three-tier windings
are commonly employed on two-pole and six-pole machines, or where the number

THE ELECTRIC CIRCXJITS 9&

of polea is not a multiple of four, Where the number of poles is a multiple of four
it is more convenient to employ the diagram shown tn Fig. 113, which enables th&
connectors to tie in only two tiere, which may be arranged in any of the methods
depicted in Fig. 113.

A two-tier winding usually occupies so much less space than a three-tier winding
that the diagram shown in Fig. 112 is preferred to the diagram in Fig. 111.

Fia. Hi. — ThiM-pbaw winding In two tlera, tbne cnUa pel group.

It will be seen that in the diagram in Fig. 112 one of the groups of coils of
phase A is long at the front end and short at the rear end. Under the next pole
a group of coils of phase C is long at the front end and short at the rear end, while
the next group of coils of phase A is short on the front end and long on the rear
end. It will thus be seen that it is only at every fourth pole that the winding
repeats itself.

Fig. 114 shows the general appearance of a winding of the type shown dia-
grammatically in Fig. 113, but with three coils per group.

96 DYNAMO-ELECTRIC MACHINERY

Where the number of poles is not a multiple of four it is still possible to
«mploy a winding diagram similar to Fig. 113(a) for the greater number of the
poles and complete the winding by means of skew coils, as shown in Fig. 115. A
skew coil is formed so as to have one half long at the front end and the other half
long at the rear end.

Figs. Ill and 112 cover the most usual cases of concentric windings. The
concentric form of winding is very commonly employed, both on alternating-
current generators and induction motors, in tboan cases in which each coil consists
of a number of turns of wire or strap. For bar windings, however, and in some
cases even for wire windings, the lattice connector is preferred.

Fig. 1 16 shows an arrangement of lattice counectors for a winding on a three-
phase machine of the hemitropic type. (See also Fig. 103.) In this case the throw
of a connector is the full pole pitch on one end and one slot short of a full throw
on the other. In Fig. 117 is shown an arrangement of lattice connectors in which
the throw is shortened. This corresponds with Fig. 104 of the single-phase case.
Although diagrams such as Fig. 116 show only one bar per slot, it is clear that each
coil lying in a pair of slot* may consist of many turns in series, and the con-
nections between successive coils can be made just as the connections are made
from turn to turn in Fig. 104. Lattice coils of this ty^ are illustrated in
Fig. 1 19. In Fig. 134 is shown a winding consisting of lattice-type coils arranged
so that each slot contains the limb of only one coil. There are thus twice as many
sloU as there are coils.

THE ELECTRIC CIRCUITS

The lattice connector, however, is more commonly used in cases where there are
^2 bars per slot or 2 coils per slot. In this case it is convenient to represent the
"bars by long and short lines on the diagram, each long line representing a bar
at the bottom of the slot and each short line representing a bar near the
mouth of the slot. The connector must always go from a long line to a short
line, as shown in Fig. 118. If there are an integral number of slots per phase per
pole, then the connections for any one phase are exactly the same in principle as
shown in Figs. 105 and 106, and the terminals would be brought out from each
phase in the same way as described with reference to these figures. In order to

Three-phase hemitropic winding with lattice end-connectora.

ascertain which three ends should be brought to the terminals and' which three
ends should be connected to the star point, it is best to choose one of the phases
— say phase C — and draw on the conductors under one pole a large arrow head,
indicating the direction in which the current will flow at a particular instant when
the current in that phase is at its maximum ; a large arrow head pointing in the
opposite direction will, of course, be drawn upon the conductors under poles of the
opposite polarity. Taking, then, that branch of phase A which lies adjacent to
phase (7, we will draw a small arrow head pointing in the same direction as the
large arrow head of phase (7, and on that branch of phase B which lies adjacent to
phase B a similar small arrow head will be drawn. These small arrow heads
indicate the current of half the maximum value which will be flowing in A and B

W.M.

98 DYNAMO-ELECTRIC MACHINERY

at the inBtant when the current in phase C is at its maximum. This will at once
be understood by reference to Fig. 1 IS. Now it is clear that we must make the
connections to the star point so that the two iialf currents from phases A and B
run together to form the full current in phase C, and it will be Been that the other

FIO. IIT.— FulJ-pLUh, four-pole, thr««-phiis« winding, with lotUcs endH^onnsctore ol
ahoTt throw. The armature oc a tODO e.v.a. turbo-generator. SOOO volts, M> cycles, 1500

iloU with one bar per slot.

three ends can be brought to terminals, the terminal of phase C providing a full
current flowing out of the machine, and the terminals of A and B providing half
currents flowing in.

Although the end connectors of a wave winding such as shown in Fig. lltf
are longer than in the lap windings shown in Figs, 117 and 120, the wave winding

THE ELEC3TRIC CIRCUITS

is generally preferred for bar-wound machines, because with it one is able to do
away with so many special connectors between groups (see Fig. 120). The
commonest method of carrying out the mechanical arrangement of the connectors
of a wave-wound machine is shown in Fig. 129. This arrangement is often
referred to as a " barrel " winding. It gives a very neat appearance, free from
unsightly connections between groups, and has great mechanical strength and
good ventilating qualities.

ABC

FiQ. 118. — Full-jpitcb, three-phase, wave winding, with lattice connectors. Two bars per
slot. The flffore shows the method of affixing large and small arrow-heads for the purpose
of finding which ends are to be starred and which ends brought to terminals.

As three-phase machines are by far the most common of all alternate-current
machines, whether generators or motors, in commercial sei*vice, it will be worth
while to consider at some length the number of slots which can be conveniently
used with any given number of poles.

In the first place (while it is possible to use almost any number of slots by
adopting certain artifices), one would usually select a number of slots per pole
which is a multiple of three, and one would prefer not to have less than six slots
per pole. Where the number of slots per pole is a multiple of three, all that is
necessary is to lay out one or other of the windings shown in Figs. Ill or 118.

Sometimes, however, we may wish to use a die in which the number of slots
per pole is not divisible by three, and sometimes even the number of slots is not
divisible by the number of poles, and it is convenient to have a chart at hand
which will enable us to say whether a convenient winding can be employed in the
particular machine in question, using a given number of slots. It may be said at

100 DYNAMO-ELECTRIC MACHINERY

the outset that, if we are prepared to introduce Blight dissymmetry into the winding,
there is hardly any number of slots which may not be used with a given number of

poles, but leaving out of account for the moment windings which involve some
dissymmetry, we may divide the symmetrical cases into five classes.

X'*. *'. ^

iHj

! ' '■ I!

THE ELECTRIC CIRCUITS 101

Glasses of three-phase armature windings.

Class A. JfTiere the number of slots per pole is divisible by 3. In these cases
one may employ any of the windings illustrated in Figs. Ill to 118, each phase
being treated as if it were a single-phase winding. These are the commonest
cases in practice.

Class B. WTiere the number of slots is one m^e or one less than a multiple of the
pole-pairs^ and is at the same time divisible by 6. Here we may employ a pure wave
winding di\dded into six groups. An example is given below. Into this class
also fall the cases where the number of slots is two more or two less than a
multiple of the number of pole-pairs. In this class of cases we can employ a
duplex wave winding.

Class C. Where the number of slots is one more or one less than a multiple of the
pole-pairs, and is at the same time divisible by 3. Here we can employ a wave
winding, divided into three groups. An example is given below.

Class D. Where the mimher of slots is a multiple of the number of pokrpairs, and
is divisible by 6. Here one can employ a wave winding with unsymmetrical end
connections, as shown in the example below.

Class E. Where the number of slots is a multiple of the number of pole-pairs, and
is divisible by 3. Here one can employ the same kind of winding as in D, with
certain limitations.

Class F. Where the number of slots is not such as to fall within any of the above
classes, one may leave certain slots unwound, and make one of the above windings
just as if the unwound slots were not there. In cases where the number of
unwound slots is a multiple of 3, it is possible to space them so that the three-
phase winding when completed is perfectly symmetrical. The main objection to
this plan is that, in the hands of an inexperienced winder, some mistake may be
made which is difficult to rectify. This class of winding is therefore not usually
employed, unless it is imperative to use a certain die in a machine which leaves
us no alternative.

We will now give examples of each of these classes.

Class A. Where the number of slots is divisible by the number of poles and
then again by 3, so that the number of slots per pole is divisible by 3. Here the
conductor diagram is perfectly simple. We may adopt the usual practice of
distinguishing the three phases respectively by thick lines, thin lines and dotted
lines, as in Fig. 110. It is clear from this diagram that any of the methods
given in Pigs. Ill, 112, 116, 117 or 118 may be employed in making the end
connections of the conductors of the different phases. Where concentric connections
are employed, as in Fig. Ill, the diagram should show the number of tiers or
ranges in which the end connections are intended to lie. When there are three
tiers they may be arranged in the method illustrated in Figs. 142 and 348 (see
page 362). Where a coil winding is employed, each coil containing several turns
per slot, the number of slots is generally chosen so as to make a winding of
Class A.

102

DYNAMO-ELECTRIC MACHINERY

Class B. The academical single re-entrant wave winding with perfectly
uniform end connections requires a number of slots, one more or one less than a
multiple of the number of pole-pairs, and if the winding is to be broken into six
symmetrical parts, the total number of conductors must be divisible by 6. Very
frequently with this type of winding there are two conductors per slot. Here is
an example.

A ten-pole machine has 66 slots, with two conductors per slot The number
of pole-pairs = 5. Now 5x13 = 65. Add 1, and we get 66. We can have a
throw of 6 on one side and a throw of 7 on the other side, giving a double throw
of 13. Starting with the top conductor in slot 1, we go to the bottom conductor
of slot 7, then to the top of slot 14, and so on until we arrive at the top of
slot 66. If we had not added 1 to our 65 we should (with a constant throw)
have arrived at slot 1, and the winding would have been closed too soon. As it
is, we pass on according to Winding Table I., and we do not close the winding
until we have passed 13 times round the machine. The last step which closes the
winding is the step from the bottom of slot 60 to the top of slot 1. This is an
example of a retrogressive winding.

Winding Table I. 66 slotB, 132 conductors, 10 poles.
Wave winding, Class B. Two bars per slot.

Top.

Bottom.

Top.

Bottom.
20

Top.

Bottom.

Top.

Bottom.

Top.

Bottom.

Now, let us divide this winding into six equal parts, each consisting of 22
conductors. Let one part begin on the top of slot 1 and end on the bottom of
slot 5. Let that part be completely disconnected from the rest. Take a second
part beginning with the top of slot 12 and ending with the bottom of slot 16,
a third with the top of slot 23 and ending with the bottom of slot 27, and so on
as indicated on the table, where the first conductor in each section is indicated by
the larger type.

Now, it is easy to see that the phase of the r.m.f. generated in the first section
of 22 conductors starting in the top of slot 1 is exactly 180'' out of phase with the

THE ELECTRIC CIRCUITS 103

fourth section of conductors starting from the top of slot 34, because slot 34
is exactly half way round the machine from slot 1, and it occupies exactly
the same position with respect to the sixth pole that slot 1 occupies with regard
to the first pole. If we therefore reverse the terminals of this fourth section of
conductors, we may connect it in series with, or in parallel with, the first section
of conductors.

The phase of the second section of conductors, starting from the top of slot 1 2,
is exactly 60 degrees removed from the first section, because when we have reached
slot 12 we have climbed through one-third of the 180 degrees between 1 and 34.
Similarly, there is a difierence of phase of 60 degrees between the E.M.F.
generated in the third series beginning with the top of slot 23 and that generated
in the second series, because when we have reached slot 23 we have climbed
through another third of the 180 degrees. It will be seen that, as there are
6*6 slots per pole, there are 2*2 slots per phase per pole. The top conductors in
slots 40 and 39 belong to the first series or phase, while the top conductors in
slots 38 and 37 belong to the second, and the top conductors of slots 36 and 35
belong to the third. In some places, however, there are three conductors lying
together which belong to the same phase, as, for instance, the top conductors of
slots 1, 66 and 65, and the bottom conductors of slots 7, 6 and 5. It is in this
way that we get the fractional number of slots per phase per pole.

As the number of conductors, 22, in each section is even, the section ends at the
same side of the machine as it begins. Thus all the connections are made on one
side. Whenever the number of conductors in a section is odd, the section ends at
the side of the machine opposite to that on which it started. It would be very
inconvenient to bring connections around the back of the yoke to put the various
parts in series or in parallel. When the number of slots is divisible by 6, and
there are two conductors in each slot, the number in each series must be even.
This is why we draw a distinction between the cases where the number of slots
is divisible by 6 and the cases where they are only divisible by 3. In the latter
case, as we shall see, a method of winding is still available without bringing
connections around the back of the frame.

WindingB with one bar per slot. The simplest way of treating wave windings
with one bar per slot is as follows : If we take any number which could be used
as the number of slots in a two-bar-per-slot wave winding as given above, that
number when multiplied by 2 gives a possible number for a one-bar-per-slot
winding. We can in this case take two slots, one odd and one even, and regard
them as one slot, the odd representing the top of the slot and the even the
bottom. Table II. gives the winding table of a one-bar-per-slot winding for
the case where there are 10 poles and 132 slots. Here the double throw is 26.
The single throw is 13 on each side. Thus we pass from an odd bar to an
even, to an odd and so on. By comparing Tables I. and II. we get a clear
idea of the relation between a two-bar-per-slot and a one-bar-per-slot winding.
Slots 1 and 2 in Table 11. just take the place of the top and bottom of slot 1
in Table I.

Now we can go a step further. Having doubled the number of slots, we
can if we like employ a duplex winding with two conductors per slot. There

104

DYNAMO-ELECTRIC MACHINERY

will be two winding diagrams. Each will be exactly the same as Table II. ^
except that the columns will be headed "Top" and "Bottom" alternately.
In one table we will begin at the top of slot 1 and go to the bottom of slot 14,
then to the top of slot 27 and so on. In the other we will begin at the bottom
of slot 1 and go to the top of slot 14 and so on. The total conductors in
each of these tables can then be broken up into six sections each of 22 con-
ductors, and the four sections of each phase thus obtained can then be combined
either in series or in parallel as desired. In Table VII. we denote this winding
by -Sg. A sample of a duplex winding is given in Table IV. The terminals of the
two windings here will be at opposite ends of the machine. This can be changed
by opening the second winding at 1 2 instead of at 1 .

Winding Table II. 132 slots, 132 conductors, 10 poles.
Wave winding, Glass B. One bar per slot.

Odd.

Eveu.

Odd.

Even.

Odd.

Even.

Odd.

Even.

Odd.

Even.

106

118

131

103

116

129

101

114

127

11*2

125

110

123

108

121

106

119

132

104

117

130

102

115

128

100

113

126

111

124

109

122

107

120

Similarly, if we multiply the 66 by 3 and have 198 slots, we can make a
triplex winding. There will be for this three tables. The first of these will
begin as follows: Top of 1 to bottom of 19, to the top of 40 and so on. The
second table will run : Top of 2 to the bottom of 20, to the top of 41 and so
on. The third vnW begin : Top of 3 to the bottom of 21, to the top of 42 and

so on.

These simple independent duplex and triplex windings must not be con-
fused with Arnold re-entrant multiplex windings which are described on page 511.

Class C. This class is distinct from Class B, because the number of slots
is not divisible by 6.*

"^It maybe mentioned here in passing that where the number of slots is divisible by 3
but not by 6, it is possible to employ windings of Class B by putting four conductors per ^ot
arranged in a double barrel winding.

THE ELECTRIC CIRCUITS

105

Suppose that we have 10 poles and 69 slots with two conductors per slot.
We see that (5 x 14)- 1 =69. A double throw of 14 will give us a progressive
winding, as shown by Table III.

Winding Table III. 69 slots, 138 conductors, 10 poles.
Wave winding, Glass C. Two bars per slot.

Top.

Bottom.

Top.

Bottom.

Top.

Bottom.

Top.

Bottom.

Top.

Bottom.

Now, if we were to divide this wave winding up into 6 equal parts of
23 conductors each, there being an odd number of conductors in each part,
it would be necessary when connecting up the various parts to carry connections
from the front to the back of the machine. This being undesirable, the following
plan may be adopted: Instead of breaking up the winding into parts of
23 conductors each, let the first part have 22 conductors, the second 24, the
third 22 and so on. Thus the first series will begin at the top of slot 1 and
end at the bottom of slot 10.

The second series will begin at the top of slot 17 and end at the bottom
of slot 40 and so on. Thus we will get the six parts. The first conductors
of each of these is indicated by the larger type. The italic type indicates
the division which would have resulted in an odd number of conductors
in each series. It will now be seen that if we connect the first part, (let us
call it phase A\ in series with the fourth part (which is also phase A)y we
shall have a series of 46 conductors, which is exactly 120 degrees of phase
removed from the series of 46 conductors obtained by connecting the third
part (beginning top of slot 47) in series with the sixth part (beginning with
slot 40). In connecting the various parts in series regard must be had to the
polarity.

It is also possible to have duplex windings belonging to Class C. An example
is given in Table IV. of an 8-pole winding in 90 slots.

106

DYNAMO-ELECTRIC MACHINERY

Winding Table IV. 90 slots, 180 conductors, 8 poles.
Duplex wave winding, Class Cj. Two bars per slot.

Top.

Bottom.

Top.

Bottom.

Top.

Bottom.

Top.

Bottom.

71.

Bottom.

Top.

Bottom.

Top.

Bottom.

Top.

Bottom.

Top.

. 81

Class D. Windings of this class are the most generally useful for low voltages,
and may be used for voltages up to 3000 on large machines. They have practically
superseded the old academic wave windings with symmetrical end connections,
except in those cases where the number of slots happens to fit the old winding.
The distinguishing feature of Class D is that the number of slots to one pair of
poles is a whole number. There may be any whole number of slots per pair
of poles: 4, 5, 6, 7, 8, or 9. The only condition is that the total number of
slots shall be divisible by 6. Thus, if we have 12 poles and 84 slots, that is
7 slots per pole or 14 per pair of poles, we can make a wave winding faUing under
Class D. If the winding is to have a full pitch, that is 7 slots, it will begin at the
top of slot 1, go to the bottom of slot 8, to the top of slot 15 and so on, as shown
in Table V. There are 13 special connectors required with this winding table, each
having a throw of 8 slots, from the bottom of 78 to the top of 2, from the bottom of
79 to the top of 3, etc. The method of breaking up the winding into six sections is
indicated as before by printing in larger type the first conductor of each section.
In this case the first section (beginning with conductor 1) can be put either in

THE ELECTRIC CIRCUITS

107

series or in parallel with the fourth (beginning with conductor 8). If there had
been an odd number of slots per pair of poles, the winding of this class would
fitill have been possible. For instance, with twelve poles we might have 13 slots
per pair of poles, with a throw of 6 on one side of the machine and 7 on the other.

Winding Table V. 84 slots, 168 conductors, 12 poles.
Wave winding, Class D. Two bars per slot.

Top.

Bottom.

Top.

Bottom.

Top.

Bottom.

Top.

Bottom.

Top.

Bottom.

Top.

Bottom.

•84

Class E. This class is the same as Class D, except that the total number of
slots is only divisible by 3 and not by 6. Here we have recourse to the same
plan of dividing up into slightly unequal sections so as to get an even number
of conductors into each section. This will be seen at once from Table VI.

Winding Table VL 75 slots, 150 conductors, 10 poles.
Wave winding. Class £. Two bars per slot.

Top.

Bottom.

Top.

Bottom.

Top.

Bottom.

Top.

Bottom.

Top.

Bottom.

108 DYNAMO-ELECTRIC MACHINERY

Class F. Where the number of slots available does not permit of any of
the above symmetrical windings, it is always possible to make an unsymmetrical
winding, and where the number of slots is great the dissymmetry can be made so
small that the divergence of the angle of phase difference between the phases
from the correct 120 degrees will not matter. These unsymmetrical windings
can be made by leaving unwound certain slots and making a winding just as
if we have the number of slots we require. The unwound slots should be
evenly distributed around the armature. Where the number of unwound slots
is divisible by 3, it is usually possible to distribute them so that there is no
deviation from the angle of 120 degrees between the phases.

The leaving of slots unwound is sometimes deliberate even when a winding
could be made using all the slots. In cases where there are very few slots per
pole, say only 3, and we are afraid of the wave-form being distorted by the
teeth, it is a good plan to depart from the winding which employs a whole
number of slots per pole. Suppose that we are designing a 30-pole machine
with a very short pole pitch, in which there is room for only three or four slots
per pole. Suppose further that to get the E.M.F. we want about 180 or 190
conductors. We would not choose 90 slots even though they are available. It
would be better to choose 96 slots and leave 6 slots unwound (see p. 305).

It is convenient to have a table such as Table VII. below, from which one
can see at a glance, whether with any given number of slots one can use any
of the windings falling under Classes A, B, C, D or E, with a particular number
of poles. Suppose that we are designing an 8-pole 'A.c. generator requiring about
180 conductors, and that we have an Annature punching with 90 slots. It is
not at first sight evident that we can make a perfectly symmetrical three-phase
8-pole winding with 90 slots and 2 conductors per slot. On referring to the
table, we see that we can have a duplex re-entrant winding of the type denoted
by Cg. Take 2 from 90 and get 88 = 8 x 11. Thus, with a single throw of 11
we will get a retrogressive winding which falls short by 2 conductors each time
around the machine. Next to this retrogressive winding there will be another
lying in the same slots, and after each has been broken up into its 6 parts,
the various parts which are nearly in phase can be connected in series with one
another. See page 104 and Table IV. It is interesting to notice in connection mth
Table VII. that each number of poles has a law of its own as to the numbers of
slots that can be used with it. This is seen from the way that the letters giving
the types of winding recur in a regular sequence in each column, each column
having its own particular sequence. For instance, the sequence in the 10-pole
column is A or D, B, (7, B^, E, Cg, C, J5, ^ or D; the only apparent exception
to this is that B2 and C^ are sometimes interchanged, but even in this there
is a law, for we have B2 when the corresponding number is divisible by 12
and Cy when it is not. Sometimes we want a winding which shall have some
of its terminals on one side of the machine and some on the other, as for
instance in the case of an A.c. booster (see page 547). In these cases we choose
a number of slots denoted by a C or an i^ ; and by keeping an equal number
of conductors in each of the 6 parts, as explained in connection with Table III.,
we will get what we desire without any dissymmetry. Very often, however.

THE ELECTRIC CIRCUITS

109

the number of slots or the number of conductors required will not permit of
this, and we will then be compelled to leave out one conductor in order to finish
at the side of the machine opposite to that on which we started. The amount
of dissymmetry introduced by this is generally of no importance.

Table VU., giving the numbers of slots that can be used with a given number of
poles to form a ssrnunetrical 3-phase winding, there being two conductors per slot.

6 Polks.

8 PoLn. 1

10 Poles.

12 Poles.

U Poles.

16 Poles.

18 Poles.

Typo of
Winding.

Type of
Winding.

■
SB

Typo of
Winding.

Type of
Winding.

Typo of
Winding.

•

1
33

Type of
Winding.

•

Typo of
Winding.

A,D

D,A

A,D

1 78

! ^

108

A,D

«0

'. 90

A,D

117

«3

126

102

136

108

A,D

<^2

144

A,D

114

105

A,D

153

120

111

102

162

A,D

126

114

106

171

«1

102

132

120

111

180

105

138

126

A,D

114

189

«7

108

144

A,D

132

120

198

A,D

111

150

138

126

207

114

166

141

129

216

A,D

120

A,D

162

147

135

226

126

168

153

138

234

102

129

174

156

144

A,D

243

105

132

180

A,D

162

150

252

108

A,D

108

135

186

168

A,D

153

261

111

138

192

174

169

270

A,D

114

141

198

180.

162

279

117

144

204

183

168

288

120

A,D

150

A,D

210

189

174

297

123

156

216

A,D

195

177

306

126

A,D

126

(^2

169

222

198

183

315

129

162

228

204

186

324

A,D

132

166

234

210

A,D'

192

A,D

135

168

240

216

198

138

171

246

222

201

141

174

252

A,D

226

207

144

A,D

144

A,D

180

A,D

258

231

210

147

186

264

237

160

189

270

163

110

DYNAMO-ELECTRIC MACHINERY

Table VII. {continued^ giving the numberB of slots that can be used with a given,
number of poles to form a symmetrical 3-phase winding, there being two
conductors per slot.

20 Poles.

22 Poles.

24 Poles.

26 Poles.

28 Poles.

SOP01.BS.

32PoLfiS.

No. of Slots.

Type of
Winding.

TVpeof
Winding.

•
•

Type of
Winding.

5z;

Type of
Winding.

i
d

Type of
Winding.

•

Type of
Winding.

•

Type of
Winding.

■B»

102

105

108

117

111

120

108

120

129

114

135

111

132

126

150

114

102

120

144

A,D

144

138

165

126

108

132

A,D

156

A,D

141

180

A,D

129

111

144

168

153

195

144

120

A,D

153

180

156

210

159

129

1 156

192

183

168

A,D

225

162

132

165

204

i D

195

180

240

174

138

1 174

216

A,D

207

183

255

117

141

177

228

210

195

270

A,D

192

A,I>

160

186

240

222

198

285

207

195

198

A,D

252

234

A,D

210

300

210

162

210

264

246

222

315

222

168

219

276

258

225

330

225

171

222

288

A,D

261

237

345

240

180

A,D

231

300

273

240

360

A,D

255

189

240

312

285

252

A,D

375

258

192

243

324

288

264

390

270

198

252

336

300

267

405

273

201

c ,

264

A,D

348

312

A,D

279

420

288

A,I>

210

276

360

A,D

324

282

435

303

219

285

372

336

294

450

A,D

306

222

288

384

339

306

465

318

228

297

396

351

309

480

321

231

306

408

363

321

495

336

240

A,D

309

420

366

324

510

249

318

432

A,D

378

336

A,D

252

330

A,D

444

390

A,D

258

342

456

261

351

270

354

279

363

282

372

288

375

291

384

300

A,D

396

A,D

•

THE ELECTRIC CIRCUITS

111

Table VIL (continued), giving the numbers of slots that can be used with a given
nnmber of poles to fonn a symmetrical S-phase winding, there being two
conductors per slot.

S4 Poles.

86 Poles.

40 Polks.

44 Por.KR.

48 Polks.

56 Polks.

64 Polks.

•

■

•

1
d

©a

•

0
d

66
84
102
120
135
138
153
171
186
204
216
222
237
246
255
273
288
306
318
324
342

B
D
B
C

E
C
B
A,D

B
C

A,D

180
198
216
234
252
270
288
306
324
342
360
378
396
414
432
450
468
486
504
522
540
558
676

A,D

a,d'

A,D

180
198
201
219
222
240
258
261
279
282
300
318
321
339
342
360
378
381
399
402
420
438
441
459
462
480

c
c

A,D

c
c

A,D

C
C

A,D

174
177
198
219
222
240
243
264
285
288
306
309
330
351
354
372
376
396
417
420
438
441
462
483
486

D
C

A,D
C

^2
^2

C
D
C

^2
^2

A,D

B
C

192
216
240
264
288
312
336
360
384
408
432
456
480
504
528
552
576
600
624
648
672
696
720
744

A,D

A, J)

A,D

195
198
222
225
252
279
282
306
309
336
363
366
390
393
420
447
450
474
477
504

^2
^2

D
0

^2
^2

A,D
C

^2
C

A,D

192
222
225
255
258
288
318
321
351
364
384
414
417
447
450
480
510
513
543
546

A,D

O2
C

^2
D

^2
C

A,D

^2
C

^2
D

^2
C

The effect of chording the winding. For the values of Kg, given on page 30,
we have assumed that we have a full-pitch star winding, that is to say, that the
conductors in series with one another are as nearly in the same phase as it is
possible to make them in a three-phase star-wound armature evenly distributed
over the armature surface. It is convenient sometimes to reduce the pitch of the
coils, either for the purpose of changing the E.M.F. or modifying the wave-form, or
it may be that for mechanical reasons we may wish to make the coil with a short
throw (see pages 113 and 114 for windings employing a short throw). We will
consider here the effect on Ke of changing the pitch of the coils.

The most simple method of finding the change which will occur in the resultant
E.M.F. when we alter the throw of the coils, or in any way alter the phase of one

112 DYNAMO-ELECTRIC MACHINERY

part of the winding with respect to the other, is by the vector summation of chords
drawn within a circle to represent the various parts of the winding. This method
may be described as follows.

We are, of course, considering alternating £.m.f.'s, and we are supposed to be
measuring the square root of the mean square values, but what is said of the
relation between these values is equally true of the relation between the maximum
values, where the wave-form is sinusoidal.

Draw a circle to represent the perimeter of the armature of a two-pole alter-
nator. Take a small arc on this circle, so short that it may be regarded as nearly
straight. Let the length of the chord of this small arc represent the sum of
.the K.M.F.'s generated in a certain number of conductors (it does not matter for

this purpose how many), which are placed near

O together (say, in oue slot), and whose e.M.f.'s are
nearly in phase with one another. Take 10 of
these short arcs, as shown in Fig. 122, so as to
obtain a larger arc whose departure from the straight
line is noticeable. The ratio between the vector
sum of all the little arcs (that is to say, the long
chord) and the length of the whole arc is equal to
the ratio between the resultant E.M.F. generated in
all the conductors lying on the arc and the arith-
metical sum of all the e.m.f.'s generated in the same

^ , ,^ conductors. The ratio of the len£i>h of the chord to

Fig. 122. °

the length of the arc is the breadth coeflScient* of
a coil occupying the arc under consideration. The breadth coefficient of a group
of coils uniformly distributed over half the circumference of a two-pole armature
is the ratio of the diameter of a circle to the half circumference, or 2-7-7r = 0'637.

Next, we must consider the effect of connecting in series a number of groups
lying in different phases.

It is necessary when using the method here described to have a very strict
convention as to sign when adding the effects of different phases. The following
convention, if carried out strictly, will avoid errors. We are to find the voltage
between two terminals of a certain machine, for instance, between the terminals
A and J? of a three-phase machine. Trace through the diagram of the winding
from A to B, and mark on our circle the arc occupied by the various sections of it,
adopting the following convention as to sign : If in tracing through from A to B
we pass from front to back on the machine in a certain section of the winding,
mark the arc on the circle which represents that section with an arrow head t
which points clockwise on the circle. For any section of the winding in which we
are passing from back to front mark the arc representing that section with an
arrow head pointing counter-clockwise.

* The breadth coefficient is ^lILf , where <r is half the angular breadth of the coil (see Fig. 321,

p. 3(J5). . <^

This factor ~ is often termed the ** winding factor."

tNote that this convention will itself take care of the question of the polarity of the poles.
Wc must, therefore, follow it strictly, never minding the polarity.

THE ELECTRIC CIRCUITS

113

After we have arrived at the terminal B, we will have on our circle a number
of arcs, each marked with an arrow head. Draw chords to all these arcs, and put
an arrow head on the chord to correspond with the arrow head on the arc. The
B.M.F. generated in the winding from A to J?, will be the vector sum of all
the chords, so taken that the arrow heads follow one another consecutively.

In the first example, we will take the straightforward case of an ordinary three-
phase generator. The result we know quite well without a vector diagram, but it
^serves to illustrate the method.

ESxAMPLE 19. Take the winding diagram of a two-pole, three-phaae, star-wound machine,
j^ven in Fig. 1)6. By way of fixing a datum line, take the centre line of one of the poles as
lying on the line EF. Then the centre line of the other pole will be on the line OH, and we
have 180 degrees of phase between them. Beginning at ^ , we trace through this phase to the
star point, and in doing so we pass from front to back thn>ugh conductors occupying a phase-
l»and 60** in width. This phase band we mark off on our circle (Fig. 124 at AiA^t and affix the

£\

FIG. 124.

Fig. 125.

olockwise arrow head. At the same time we have passed from back to front through con-
ductors occupying a similar phase band in front of the opposite pole. This we mark off on our
<sircle at .^^^4, affixing a counter-clockwise arrow head. Passing on again from the star point,
we go through phase B, some of the conductors being traversed from front to back. These are
marked off on the arc ^1^3 with a clockwise arrow head. And some of the conductors are
traversed from back to front in the position of the arc ^3^4, and are marked with a counter-
•clockwise index. Now, the resultant e.m.f. generated in these conductors is proportional to
the length of the vector A^B^, which is built up of the vectors A^A^, A^A^, B^B^, ^3^4 (s^e
Fig. 125). In this case the vector AiB^ is 0*866 of the arithmetical sum of the small vectors.

In the second example, we will take the case of a two-pole, three-phase winding
having a very short throw.

ExAMFTJB 20. In Fig. 126 is given the diagram of a two-pole, three-phase winding, lying in
-24 slots. The throw of the coils is very short, being just a little over 90°. There are supposed
to be two paths in parallel, but for the s«ike of simplicity the end connectors are only shown on
one of the paths. The end-connector diagram for the other path is exactly the same as that
shown, except that it is pushed forward 12 slots, or 180". To find the ratio of r.m.f. generated
in this winding to the K.M.F. generated in 4 conductors lying in adjacent slots, we describe our
circle as before. In Fig. 127 we have made small circles to note the position of the slots for
ease in following the diagram, but this is really unnecessary. We wish to find the E.M.F. at
the terminals A and B. Imagine that the centre of one of the poles is, for the instant, opposite
w.M. H

114

DYNAMO-ELECTRIC MACHINERY

the datum line EF. This gives us the phase position of the vertical line EF (Fig. 127).
Trace the winding tlirough from terminal A to terminal B. We pass from front to back.

-A short-chorded, three-phase winding for a two-pole turbo-generator. Throw

of coils two-thirds of pole pitch.

through slots 1, 2, 3, 4, and therefore mark a cliord (in Fig. 127) of tlie arc which embracea
1 and 4 with a clockwise arrow head. We pass from back to front through slots 8, 9, 10, 11»
and therefore mark the corresponding chord (Fig. 127) with a counter-clockwise arrow head.

Fio. 127. Fio. 128.

Showing method of finding the voltage generated in a short-chorded winding.

This leads us to the star point, from whence we pass into phase B, which leads us from front to
back through slots 16, 17, IH, 19, as denoted by the clockwise arrow head in Fig. 127» and from
back to front in slots 9, 10, 11, 12, giving us the counter-clockwise arrowhead. This brings us

THE ELECTRIC CIRCIJITS 115

to terminal B. We now make u summation of the veotora, as shown in Fig. 128. Th« ratio
of the length of the vector 1, IB to the vector 1, 4 ia the ratio wo require. It will be seen thM
in this case the K.M.r. generated in the 16 conductors in series in Fig. 126 is only O'S of the
■.H.P. of the 16 oondnctora of a similar machine connected as in Fig. 116, as seen from a ooin-
puiaon of the vectors A,Bi (Fig. 125) and 1, l9(Fig. 128). If. therefore, we had a three-phfMB
windiog with S oouduotors per phase, which, when conneoted as in Fig. 1 16, gave us S50 rolts,
weoonkl, bf ohording the winding ns shown in Fig. 128, obtain R50x 0*06 = 440 volt«.

It is very oecessary to follow strictly the convention aa to the arrow heads od
the dH^ram, as the vector |)olygon for some chorded windingB will contain vectors
which are almost directly opposed to one another, and it is impossible to be sure of
the value of the sum unless we have taken strict account of the true phase position
of all the oomponente.

Pia. IHfl. — Barrel end'Connscton of thm-phaie siuumtor anDatuni wound with two
ban p«r slot and EODA«cC«d, u shown In Fig. IIS, to nneratfl 440 ftAtt. The ban Id this
osM ace bent after being pot Into the ssml-oloHd slots.

if»ti antral amnfement of windings. The mechanical arrangement of the
winding will depend upon the type of end connections we have chosen. End

116 DYNAMO-ELECTRIC MACHINERY

connections of the lattice type permit of a very simple mechanical arrangement.
In general, they will form two layers. If these layers lie on a cylindrical or
conical surface, we have what is commonly called a " barrel " winding. Such a
winding, consisting of solid bars, is illustrated in Fig. 129. In this case there
are two bars per slot, one of which is bent to the right and the other to the

ig of itMOT supported by li

Fid. 130a. — Bunl winding showlni metliod ol (lilng to liuiiUt»d liogi Bapportfld on bracksU.

left to form the lattice work. Fig. 130 and 130o show methods of fixing h
winding of this kind when used on turbo-generators. Fig. 13a shows a somewhat
similar method of clamping.

A barrel winding is frequently built up with wire-wound coils. Such a winding
on the atator ol an induction motor is shown in Fig. 119.

Where such a winding forme part of the revolving element of the machine, it
is usual to support it mechanically by means of a wire band or end bell. Fig. 131
shows an ordinary c.c. armature with barrel winding. Fig. 133 shows a similar
winding on the rotating field-magnet for a turbo-generator. The advantages of

THE ELECTRIC CIRCUITS

C. umatnre reidr for b

¥ia. ISE;— "Shart"-t;p« wliidiiis oi

118 DYNAMO-ELECTBIC MACHINERY

this type of winding are that it can be easily formed iuto shape, easily insuUted in
a satisfactory manner, and where the radial thickness of the winding is not too
great, the cooling is very good.

Where the throw of the coils is great, as in two-pole niachines, this type of
winding projects rather a long way from the armature iron, and therefore takes

more copper than some of the other types. In cases where there is very little end
room, as in railway motors, this winding is modified to form the " short " type shown
in Fig. 132 (see page 163), Where the throw of the coils is short, as in machines
having six poles, or a greater number of poles, the "barrel" winding is very
economical in material.

On stationary armatures the "barrel" winding is sometimes formed from coils
arranged bo that there is only one limb of a coil per slot. Such a winding is
illustrated in Fig. 134.

THE ELECTRIC CIRCUITS 119

The "barrel" winding will be found useful in coees where it is desired to make
<4e].throw of the coils much shorter than the pitch and interleave the different
phases. In this case all coils must be completely insulated to sUnd full voltage to
-earth and between phases.

Another arrangement of lattice end connectors employed in stator windings is to
bend up the conductors until they lie at 30° or 45° to the axis of rotation. This
■winding can be secured by means of suitable bracket*. Fig. 135 shows the
annature winding of a 10,000 K.V.A., three-phase, 2400-volt, 60-cycle, four-pole
^toerator built by the Westinghouso Electric & Manufacturing Co. of America,

OompBDy).

The coils have a throw of only two-thirds of the pole pitch, so as to make the
end connectors shorter (see Fig. 126). Thus each phase of the winding occupies
a coil-breadth of 120 electrical degrees, and this has the effect of eliminating the
action of any third harmonic even if the phases are connected in mesh (see page
307). As coils belonging to different phases lie in the same slots,* each complete
coil must be insulated all over to withstand full pressure to ground. Annatures

•With a ohorded winding, the eddy-ourrenta {see p. 144) in an Bmmliire eoiiduulor may bo
smaller than with a fnll-piteh winding. To calculate tlie eddy-ourrent loss, we van ub« the
carves givBo in Figs. 167 and 167a, but we (Ake f nictioiuil valnes for m. A. B. Field has given
Cbe lollowinK valaes for m for the case where there are two conductors per slot, tliu current in
tbe two conductors being out of phase. For thn inner conductor ir= 1. For tlia outer oon.
dnctorwe have ni = 2 for zero phase differBnoB; I 82 lur 60° difference ; 142 for QO" difference ;
and 1-0 tor ISO" difference. For o threo-phoee armature Bhorl-ohorded by 60°, the eddy-current
kMN generally comee out some 76 per cent, of the value of a non-chorded winding.

120 DYNAMO-ELECTRIC MACHINERY

of tbia type have mlhetood repeated abort circuite. The spaces between a>ils
permit of excellent ventilation. If we bend up the conductors still more, until
they lie in h plane at right angles to the Kxis of rotation, we have what is
commonly called an involute winding. Such a winding is illustrated in Fife.
136 and 137.

The advantage of this arrangement is, that it enables a winding having a
long throw to be made without extending the length of the armature too much

., three-phue, 24a0-volt. W-cydi, tour-pole gnnentor,
ice winding it vi uigle ol i5° to the aiU.

in the direction of the shaft. It can be convetii<;ntly clamped against the flat face
of the end plate, but requires considerable radial depth where the throw is great.
Such an arrangement would result from any of the lattice diagrams given in
Figs. IO;t, 105, 107 or 110. Where, however, the end-connector diagram is like
Fig. 104, the involute does not show a continuous lattice work. Fig. 117 shows
iin involute winding with short-throw end comiectors. This winding can also l>e

THE ELECTRIC CIRCUITS 121

supported mechanically by means of clamps fixed by bolts passing through the
openiuga in the ninding.

FiO. 136.— Involute gUtor vrlndiD(.

For bAQd-wotmd induction motors with semi-closed slots, the type of winding
illustrated in Fig. 138 is commonly used. This is analogous to an involute
winding, but the coils are merely formed of cotton-
covered wire wound promiscuously and bent into a
skew shape so as to clear one another. Coils of this
type can be put turn by turn into semi-closed slots
which have been previously insulated, the mouth of the
slot being subsequently closed by wooden or paper
wedges. This is sometimes called a "vutsh" winding.
In Fig. 139 is shown a "mush" winding on the rotor
of an induction motor. In Fig. 140 the first inaei-ted
coils are seen forced back so as to allow the last eoils
to be put in.

Concentiic coils. Where each coil consists of a
number of turns and it is required to insulate the
phases very well from one another, some makers prefer
to use concentric coils so as to preserve larger insulating
spaces between coils belonging to difTerent phases than
is generally possible with the lattice type of end con-
nector. If the slots are closed or semi-closed, the
concentric coib are sometimes wound by hand through
insulating tubes, the end connections being subsequently
insulated. Hand winding cannot be recommended where

the voltage is high. Wire-wound coils are much more d™p o^nlL^M^JjSS'ng'
satisfactory when fonned and impregnated before being

placed in the slots. Open slots are therefore desirable where the number i
conductors per slot is great.

122 DYNAMO-ELECTRIC MACHINERY

Where the number of conductors per slot ia few — say one to six — a very

satisfactory method is to make the

part of the coil which lies in the
slot of straight conductors insulated,
impregnated and wrapped. Aft«r
these have been placed in the slot
(which in this case can be made
closed or semi-closed), the end con-
nectors, which have likewise been
previously insulated, can be jointed
to the slot conductors, each joint
being thoroughly insulated as it is
made. The advantage of this
method of winding is that it enables
the conductors which lie in the slot
to preserve their perfect straightness
throughout the whole insulation
ti-eatment, and a more satisfactory
coil, free from air spaces and bulgy
insulation, can be made than where
_ „ , „ , a large coil is insulated as a whole.

motoi, put wire br wire ttuousb Uie moutlu o[ aemi-doKd Moreover, in case of a breakdown il

alotiuid labsequently iOBUtaMd. . -, ,

IS possible to remove any one coil
of this type without disturbing the remainder. The joints between the straight
conductors and the end connectors will not cause any difficulty if they are not too
numerous, and if plenty of apace is
aUowed for jointing and insulation.
A concentric winding of this type,
made in two tiers, is illustrated in
Fig. 141 (compare Fig. 112). The
method of making the joints is seen
in Fig. 142, which shows a winding
in three tiers (compare Fig. 111).
For singlepluise iniiditttff, such as
illustrated in Fig. 102, a concen-
tric coil is very suitable, and ofTers
no special difficulties. For large
machines, however, where the span
of the coil is great, they are some-
times bent back against the end
plate, because in this position they
can be more securely fixed by means
of clamps.

Where the number of poles is

not a multiple of four, it is possible F"0- ISB.— "Mo«h" winding on Ow rotor ol indw-

■luv •• 1- , ,™— . ^j^^ motor, put wtro b» wire throngh the moatlii o( Mmi-

to employ a two-tier winding by cio»ed »ltit» md »nb«qa*iiUy iii*ui»ied.

THE ELECTRIC CIRCUITS 123

usJDg on the one pair of poles two ekew coils, as illustrated in Fig. 115. The
three-tier winding is the moat convenient to employ with concentric coils on
two-pole and six-pole machines.

Tbe forces which come into play when the windiiig of an electiic geaeiatm'
is short circuited or when an unexcited armature is thrown aaddenly cm to a
high-voltage inain. In discussing the mechanical arrangement of armature
windings it is necessary to say something about the enonnous forces which come
into play when a. winding is short circuited. Any generator or motor is liable
to this accident, and it should be so constructed tliat it will not be seriously
injured if the accident should occur.
The designer must be able to say
what windings require special
bracing, and what windings are
sufficiently strong without bracing.

In general there are two kinds
of accidents to consider. First, the
case where a generator is running
fully excited, and is short circuited
at or near its terminals, and secondly,
the case where a machine standing
idle is inadvertently switched on to
the supply mains. In the first case
we have, before the short circuit,
a high E.M.F. within the winding
which takes an appreciable interval
of time to fall to a low value. The
strength of the current depends
upon the characteristics of the

machine itself. In the second case, ^::oi!!,'?';?o™u;eSS5o!"^uSln?£'5i^^^^
there is no E.M.F. on the winding

before the switch is closed, and the length of time that the E.M.F. is exerted
depends upon the characteristics of the generators supplying the mains and the
operation of any cut-out devices which may be in circuit.

We will take first the case of a short-circuited alternate-current generator.

If the armature of an ordinary alternator is short circuited while the machine
is at rest, and the machine is then run up to speed and fully excited, the current
in the armature will not in general rise to more than 2J or 3 times its full-load
value. This is because the current lags almost 90 degrees behind the phase
position of pole centre and it demagnetizes the poles (see page 282). This current
would not be sufficient to bring into play any serious forces on the winding. But
if an alternator, while running at full speed and fully excited, is suddenly short
circaited at the terminals of the armature, the current may rise to more than
20 times its full-load value. The first rush of current is propelled by the
full E.H.F. of the generator, because there is not sufficient time during the first
«ycle after the short circuit for the field magnet to be demagnetized to any con-
siderable extent. The rate at which the current rises is determined by the

124 DYNAMO-ELECTRIC MACHINERY

Belf-inductioa of the armature winding. In coDeidering the armature self-induction
which ie effective immediately after a short circuit, we must take only that
part which is due to the flux leaking across the slots, along the air-gap and
around the end windings.

There cannot be an instantaneous weakening of the field-magnet, because, as
the current in the armature rises, there is an eddy current in the pole face, or

FIQ. 141. — Two-tier concrnttlc windlnE of 0000 e.t.j. three-phaH gentrMor (Britisfa
Weatln^iouH Compiny).

in the field winding itself, which maintains the flux from the pole almost at
its full value. It is only as this eddy current dies down that the armature
demagnetizes the field.

The eddy current in the pole or in the field winding (or it may be in both), while
exerting a magnetomotive force to keep the main flux of the pole in existence
notwithstanding the demagnetizing effect of the armature current, sets up around
itself a leakage flux in the fleld-magnet, and this leakage flux opposes the rise of
the eddy current and makes the rate of rise smaller than it otherwise would be.

In order that we may have a rough picture of what is happening, let us take
a particular case, the case where a short-circuit occurs in phase .^ of a turbo-

THE ELECTTEIC CIRCUITS 125

generator whoae field-magnet is made of solid steel of the type shown in Pig. 350.
We may, in order to fix our ideas, consider actual values that one finds in practice.
Let Fig. 150 represent a three-phase 5500-K.w. generator with four solid salient
poles revolving at 1000 ILP.U. The total flux per pole amouiite to 78,000 kilolines ;
there are 27 conductors per phase per pole, and the instantaneous value of the
E.M.F. generated in phase A at the moment when the pole is in the position
shown in Fig. 150 ie 9000 volts. The resistance of phase J is 0056 ohm. The
magnetic flux Ij, which leaks across the slots and around the end windings of

Fl9, 14i— Concentric i

the machine when 1 ampere passes in phase A amounts to 62 kilolines per pole.
For full-load current — 28S amperes ( = 405 amps, maximum) — the leakage there-
fore amounts to 2500 kilolines, or 3*2% of the total flux per pole.

In order to simplify matters, the Hgures given here for the flux per pole are
reduced to allow for the breadth coefficient, and the leakage flux is dealt with
in the same way. Thus, if jV= effective flux per pole,

volts per phase =23r«-jVx 10"", where ^ = the turns per phase.

Bftte of tIm of tho current. Now, let phase A be short circuited at the
moment when the pole is in the position shown in Fig. 150. There is an E.M.F. of
9000 volte tending to drive current through phase A. As the current rises it

126

DYNAMO-ELECTRIC MACHINERY

will not only set up leakage Aj, but it will create a magnetomotive force in all
such paths as mm, threading through the path of the iron pole. As soon as
any flux begins to grow in the path mm, it immediately produces a current in
the pole face of an amount almost equal to the total current in the phase band A
and opposite to it in direction. In the case in Fig. 150, with counter-clockwise
rotation of the N pole, the current in the phase band A will flow towards the
observer from the paper ; the eddy current in the pole will flow from the observer
towards the paper. The return path for this eddy current will be along the
sides of the pole and back along the face of an S pole. This current opposes
the creation of flux along the path mm, so that the flux cannot grow at a
greater rate than is just sufficient to generate the eddy current against the
opposition of the resistance and self-induction of its path. The self-induction of

Wl«

PhAM A

PhaeeVr

Fia. 150. FlO. 151.

Showing eddy currents in the pole-face and the paths of the leakage flux Ai and A,.

the eddy-curi'ent path is caused by magnetic leakage, which may be represented
by the symbol A^. As the flux along mm grows, it will set up a back E.M.F.
in phase A, so that the effect will be just as if the resistance and self-induction
of the eddy-current path in the pole were transferred to phase A. The value
of the resistance and coefficient of self-induction of the eddy-current path in
the pole would be difficult to calculate in any particular case, but in the case
considered below, the coefficient of self-induction (the more important term),
when multiplied by the ratio of transformation, appears from the result of experi-
ment to have a value equal to about 2'4 times the coefficient of self-induction
of that part of the armature winding which lies opposite the pole. We therefore
have the leakage \ + A.^ as the main controlling factors in determining the rate
at which the current / in phase A begins to rise; taking Aj4-Aj in the above
case to be 21 '2 kilolines per ampere, the rate at which the current would begin
to rise is 800,000 amperes per second.

If the short-circuit occurs in A at the instant when the poles are in the position
shown in Fig. 150, the current 7,,^ cannot rise higher than such a value as will make

THE ELECTTRIC CIRCUITS

127

the leakage flux equal to the working flux per pole, that is, /rt(^i + ^) = ^» K>
however, the short-circuit occurs when the poles are in the position shown in
Fig. 151, the total change of flux threading through phase ^, as a North pole is
replaced by a South pole, is 2N; so that the limiting value of the short-circuit
current is such that Ia(^i + A.^) = 2N.

Fig. 152 shows generally the way that the short-circuit current would rise, if
we leave out of account the effects of resistance and capacity. The height to
which it will rise depends upon the instant at which the short circuit occurs. If
the short circuit occurs when the voltage is at its maximum, the current begins

aooo

... ^

/
/

9000

TOOC

900C

'/^

•

0OCO

*ooe

'/■

Zero k

ine c/

Short

arcvtt

Currm

SOOO

/
/

dep***^* or* M# msfant «/ m^hich
ShOff ctrct/it occurs , mg for thm

meee

/
/
/
/
/

throv\

tr -oo*

% aro
OO A^

» mno

/ooo

1
f

/
1

JyjLk

oad C

urr*n

— s

*oa«

^ss^

■009

0«

'0*»

•om

lUf

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''9M

Fig. 152. — The magnitude and phase of the instantaneous short-circuit current as comiutred

with full-load current lagging 90°.

to rise at its maximum rate, but it cannot rise for longer than one quarter of a
period. We therefore must in that case draw the zero line for the current curve
through the ordinate 4600. in Fig. 152, and get a current curve symmetrically
placed with regard to the zero line, the maximum being not more than 4600
amperes. If, however, the short circuit occurs at the instant when the voltage
in phase A is at zero and becoming positive, then the current will have twice as
long a time in which to rise, and will rise to nearly twice the value, or nearly
9200 amperes. It does not rise to quite twice the value, by reason of the fact
that the pole is being gradually demagnetized, and the resistance of the armature
begins to have an appreciable effect at the higher values of the current.

As the pole moves from the position shown in Fig. 150, the eddy current
moves along the pole face as though it were a reversed reflection of the current in

128 DYNAM0-ELEC5TRIC MACHINERY

the phase band A mirrored in the face of the pole, and shortly after the comer of
the pole has left phase A^ there is an eddy current in the pole face down one
side and up the other, which tends to keep up the flux in the pole against the
strong demagnetizing current in phase A, which is now in the best demagnetizing
position (see Fig. 151).

If instead of a solid pole we have a laminated pole, the resistance of the path
for the eddy current is very much higher, but it is found that the exciting current
in the winding is increased at the instant of short circuit, and takes the place of
the eddy current in keeping the value of the pole flux nearly constant for the
instant after short circuit. By the time that the pole S comes under phase Ay
the eddy currents are beginning to diminish, so that pole S is not as strong as
pole N was. The current falls under the in^uence of pole S^ and rises again
under the influence of the next N pole, so falling and rising, and describing a
waving line which keeps its mean value above the zero line for several alter-
nations. It may be said that at the instant of short circuit a direct current
is generated in the winding which is maintained by the self-induction of the
winding and slowly killed by the resistance. Superimposed on this direct current
is an alternating current generated by the passing of the poles alternating N
and S, and at the same time gradually growing weaker as the eddy current in the
pole face or in the exciting windings grows weaker. After a few seconds the
poles become normally excited, and the current sinks to 2^ or 3 times its full-
load value.

There is a difficulty in dealing with this subject analytically, because the eddy
paths in the poles are not of simple form, and their resistance and self-induction
vary as the eddy current takes up different positions in the poles. We know,
however, that there is produced at the instant of short circuit a large magnetizing
eddy current. We also know that .the magnetizing eddy tends to live by reason of
the self-induction of its path, but is slowly killed by the resistance of the path, so
that we may assume tliat it changes with time t' approximately as the expression

/«€ ^ , where /« is its maximum value and r^ and l^ are the values of the resistance
and coefficient of self-induction of the path, though these are not necessarily
capable of being expressed by constants. Now the e.m.f. in the armatiu^ at any
instant may be regarded as consisting of two parts, one part e«, the E.M.F. which
is sufficient to drive the final short-circuit current through the armature against its
resistance and self-induction, and the evanescent part e^, the E.M.F. generated by
the flux from the pole which is produced by the eddy current in the pole.

Here we have written t' = {t-t^ in order to have only one time variable, t^ is
the time at which the switch is closed.

At the instant of short circuit (t-t^ — 0 and c« + ec = ^> the full E.M.F. of
the machine. As the eddy current in the pole dies out^ e^ disappears and « = e«,
and the machine then gives its normal short-circuit current. The expression

THE ELECTTRIC CIRCUITS

129

for the current at any instant after a short circuit therefore takes the form

£, + € '«

-?(«-«!)

(sin (2jm< - o)} -

£, + Ee

«-?<'-'>)

(sin '27rnt^ - a).

The last term of this expression corresponds to the evanescent term which
always appears in the expression for the current after switching on. As ^j is a
constant, the last term represents a current"^, which is always on the same side
of zero, possibly very great at the instant of switching on, and slowly dying as it
is killed by the resistance of the armature winding. .The angle a of course equals

tan-i

In the above expression r^ and \ are the apparent resistance and coefficient
of self-induction of the armature winding. We say " apparent," because if there
are any circuits (such as eddy-current paths in surrounding iron) which carry
currents induced by the armature current^ the resistance and self-induction of

* It may make the matter clearer to some readers who are not very familiar with switching
phenomena if we consider the current which flows in a single-phase circuit whose resistance is
M and inductance L when we suddenly switch on a voltage following the law : j^j = J^Qsinjo^.

The instantaneous value of the current

- —ft- 1 )
i=lQ sin (;rf - ^) - e L /© sin (pt^ - ^),

where lQ=—r==J==, <p=tan~^^, and <,— the time of closing the switch.

The wave-form i in Fig. 153 shows the values of i when t^ is just a little less than half a
period, and where L is great compared with R.

FIG. 153. — Wave-form of current suddenly switched on.

A simple way of arriving at such curves as these is as follows : We know that, finally, after
A number of cycles the current will settle down to the value

i=/Qsin(pi-^).

Plot this wave-form as shown at /. Now we know that the current at the instant of
switching must be zero. In order that it may be zero, there is superimposed upon / a uni-

directional current - /osin (p<, - 0)e -^ , which is shown plotted and marked //. This is

equal to /gsin ipt-^) at the instant t^, and beine subtracted from it makes the current zero.
It has the effect of displacing the zero line of / by an amount that is always decreasing

l>ccause of the factor € ^

In the formula ffiven at the top of this page, we are concerned with two evanescent factors :
one of these oontrolB the unidirectional current which is superimposed as in Fig. 153 because of
the sudden switohins on, and the other controls the rate of decrease of the maximum voltage
generated because of the dying eddy current in the field-magnet.

W.M. 1

130 DYNAMO-ELECTRIC MACHINERY

these circuits must (after multiplying by the proper ratio of transformation) be
added to the true resistance and self-induction of the armature, just as with a
transformer the resistance and self-induction of the secondary of a transformer
are transferred to the primary.*

In order to ascertain the instantaneous value of the current in the armature at
the time of short circuit some records were taken on a Duddell oscillograph fitted
with a cinematograph film. Fig. 154 shows one of these records taken on a
machine of the type shown in Fig. 150. The curve ^ shows the voltage before
the short circuit, which, in this case, is 3900 virtual, the maximum point being
5700 volts. At the instant of short circuit the voltage at the terminals of the
machine falls to zero, and the current curve springs into being ; the highest current
recorded was 3100 amps. The curve is sufficient to show the general nature of
the current on short circuit. It will be seen, as indicated in Fig. 151, that it rises
to such a high value during the first half of a cycle that the pole which passes
during the next half cycle is just sufficient to bring it to zero. It then rises and
falls in waves of gradually diminishing amplitude, until after a lapse of several

/\ r\,r\.r\^r\,r\^

• Voitd -scale i x cm - n,ooo volts . Current scale *, x cm . «X(v>oo ampere*
SwiCch'cloaed "^"^^ ^^'®; ao cm. - 1 second.

V- voltage measured at termmals of the g^enerator." To ^t the right phase of voltage m the ley of the star

nvtr^e the polarity and subtract so**

Fio. 154. — Oscillograph record of armature cnrrent when an alternator is short circuited.
Volts before snort circuit =3900 virtual. Maximum current 3100 amperes.

seconds it assumes the value it would have had if the short circuit had been made
before the field was excited. The value of the current at the first peak is, in this
case, 7*6 times as high as the maximum value of the current after it has settled
down.

It will thus be seen that the amount of current which flows when a winding is
short circuited is determined by voltage on the winding and the coefficient of self-
induction of the winding, and in calculating the self-induction we must take into
account only those magnetic paths around the winding through which magnetic
lines can pass without setting up opposing eddy currents. On page 422 we have
given a simple method of calculating the leakage flux across a slot, and on page 425
we give rules for roughly estimating the leakage flux around end windings. If the
sum of these fluxes for one pole in stator and rotor, when one ampere is passing in

* The foUowitiK are the values of the different quantities as caloulated from the oscillograph
curves taken on snort circuiting the 5500-K.w. generator above referred to:

^«=975 volts maximum in one phase.

^,=2740

ri =0-105 ohm.

/, =0-011 henry.

jt If i»

♦•fl-

J-" =2*06 during first half -second, but changes slowly to 1-15.

n =33 J.

tj^ =0*002; t is taken so that the volts of one phase of the generator =3715 sin 2m/.

THE ELECTRIC CIRCUITS 131

the armature, is taken, and denoted by /, and l^, then the highest possible value of

the short-circuit current is approximately Ia= ,' i ■ when A' is the flux from one

pole. We multiply by 3 on account of the doubling effect which may occur at
some instants of switehing.

The experiments are more fully described by the author in a paper,* reference to
which is given below. The reader is referred to this paper for other curves and
deductions therefrom.

Some very interesting oscillograph curves taken during the short circuiting
of a 10,000 K.V.A., 2400-volt, 60-cyele, 4-pole generator, are given by A. B, Field.t

t\a. IGSfr.-- Slngls-phiM •hoct circuit ocRurring when t!:e voltaga Is neir Its Duxlmum.

Two of these are reproduced here. Fig. 156a shows the short circuit at an instant
when the voltage was near zero, and Fig. 1656 shows the short circuit at an instant
when the voltage was near its maximum. In the first case the middle value of the
current curve is very much displaced from the current zero line, and in the second

* " Short Circuiting of Large Electric Generfttora and the lUauIttng Forces on Armature
Windings," Jcmr. Iitfl. Kite, Engrt., vol. 46, page 295.

t "Operating CharacteriBtica of Largo Turbo-Cipnerators," Amtr. la^. Elec. Eiufrt., vol. 31,
pogeOM.

132 DYNAMO-ELECTTRIC MACHINERY

case it is very little displaced, as we might expect from the foregoing theory. In
both cases the maximum change of current in a half-cycle, that is, the amount
measured from the top of the positive peak to the bottom of the negative peak,
is the same. It is, in fact, the current which could produce a leakage flux nearly
equal to twice the pole flux.

In Fig. 154 a curve has been drawn through the crests of the positive waves,
and similarly one through the crests of the negative waves, and these have been
extended back to the axis drawn for the instant of short circuit. The two curves
intercept this axis, marking off a length PQ corresponding to 37,000 amperes.
For a given machine the length of this intercept is almost independent of the par-
ticular instant at which the short circuit occurs. It is therefore useful as a charac-
teristic of the machine. It may be taken as roughly proportional to the flux per
pole, and inversely proportional to the leakage flux per ampere in the armature. It
also depends upon the number of phases short circuited. The curves given in Fig.
154 (b) and (c) relate to a single-phase short circuit at half normal voltage. The
lower curves in these figures show how the exciting current varied. The following
interesting data are given of the machine in question. With the rotor removed
and an external source of 60-cycle current applied to the stator terminals, the
impedance was foimd to be such as to give approximately 8*4 times normal current
with full three-phase voltage applied, and 7 '3 times normal current when the rated
voltage was applied across two only of the three terminals. Similar tests made
on this machine with the rotor in place indicated an impedance which was not
strictly independent of the magnitude of the current, but which apparently would
give about 12 times normal current with three-phase full voltage, and about
lOJ times normal current single phase. The rotor was of the solid steel tjrpe like
that depicted in Fig. 350, and the stator is shown in Fig. 135. The air-gap was
I inch at each side, and the stator slots 0*86 inch wide. The power absorbed on
the impedance test with the rotor in place amounted to 340 K.vv., and less than
one-sixth of this when the rotor was removed. This increase in power was, of course,
due to the heavy eddy currents in the face of the stationary rotor.

Changes of proportions which improve the regulation of the generator do not
of necessity cause an increase in the momentary short-circuit current. In par-
ticular, it should be noted that a high ratio between the ampere-turns on the
field magnet and the ampere-turns on the armature (a feature which gives good
current regulation) does not in itself increase the short-circuit current. The short
circuit is kept down by decreasing the flux per pole or by increasing the armature
leakage per ampere (see page 388).

Switcliing in when out of step. If two similar alternators running at full speed
and fully excited are thrown in circuit with one another when directly out of
step, the currrent which flows through the armature is the same as if each
machine had been short circuited at its terminals. The E.M.F. taken in the
whole circuit of the two machines is doubled and the resistance and self-induction
are also doubled. If the machines are thrown into circuit only partly out of
step, the current which circulates is not so great, being equal to the short-circuit
current multiplied by the sine of half the angle of phase displacement between
the two machines. Where two machines are feeding a busbar and a third similar

THE ELECTTRIC CTRCUITS 133

machine is thrown on to the busbar directly out of step, the current flowing in
the latter machine will be one-third greater than if it were short circuited at its
terminals. This is because the total e.m.f. in circuit is doubled, while the
resistance and self-induction of the circuit are only increased in the ratio of 3 : 2.
Where three machines are feeding a busbar and a fourth is thrown on to the
busbar directly out of step, the current is 50 per cent, greater than if the machine
were short circuited at its terminals; and so on, as the resistance and self-
induction of the machines in circuit with the busbar become less and less. The
maximum effect will be obtained where a machine is switched on to a busbar fed
by a very large number of generators. In this case the current might rise to
almost double the value it would attain on a dead short circuit ; i.e, the forces
which would come into play would be nearly four times as great.

Next consider the case of a generator or motor switched suddenly on to the
line. The currents through the armature winding in these cases may be even
greater than where the E.M.F. is generated in the machine itself. We may,
for instance, have a very large power station, the voltage of which is very little
affected by the drawing of a large current. Where the self-induction of the
windings of the generator or motor is fairly high, the current will be limited
by this circumstance. If, for instance, an engine-tjrpe generator has laminated
poles and the field circuit is open, the current which will flow through the
armature or switching on will not be as great as if the poles are solid or if the
field circuit is closed. If an induction motor has a squirrel-cage rotor, the
current flowing on switching the stationary machine on to the busbars is much
greater than if the rotor is of the wound type and is open circuited. The
leakage flux of induction motors is usually such a large percentage of the total
working flux that, even with a squirrel-cage rotor, the current on switching
on a dead machine is not sufficient to cause serious trouble. The question
which the designer must ask himself is: **What is the value of the self-
induction of the winding which would be operative in cutting down a current
on switching on, and what are the chances of the machine being switched on
under circumstances likely to injure the winding*?" Then a rough calculation
of the forces upon the winding will tell him whether it is necessary to
specially brace the winding or not, and what kind of bracing ought to be
employed.

Forces on the windini^s. The study of the instantaneous currents which flow
when an alternator is short circuited is important on account of the great forces
which they bring into play upon the armature windings. In fact, there has been
considerable difficulty in the past in devising adequate means of supporting the
coils. Even with slow-speed generators, it was known that the sudden rush of
current which occurred on short circuit, or when the generator was thrown on the
busbars badly out of phase, would injure the winding unless it were made very
strong and suitably supported. But it was not until after many serious accidents
that the designer realized how many times greater were the forces he had to deal
with in the case of turbo-generators. The reasons for the difference in this respect
between slow-speed machines and turbo-generators are as follows: In the first
place, the high-speed machine has comparatively few poles, and therefore the

134 DYNAMO-ELECTRIC MACHINERY

ampere-turns per pole are very much greater. For instance, a 3000-K.w. 25-cycle
engine-type generator, running at 94 revs, per minute, may have about 2000
ampere-turns per phase per pole, while a 3000-K.w. 25-cycle turbo-generator,
running at 1500 revs, per minute, may have as many as 8000 ampere-turns
per phase per pole. The force is proportional to the square of the current,
so that four times as many ampere-turns will give very many times as much
force. A case is worked out in the author's paper already referred to.

Secondly, the span of the coils in the engine-type machine is very much shorter.
In the machines compared above, the spans of the coils might be 18 in. in the one
case and 90 in. in the other.

Then, again, the magnetic flux leaking across the slots and around the end
windings bears a much smaller ratio to the total flux per pole in the case of many
turbo-generators than it does in the case of engine-type machines. This con-
sideration, as we have seen above, is one that determines the value of the current
on short circuit.

The troubles that were experienced in the early turbo-generators were, no
doubt, partly due to the disinclination on the part of the designer to bring the
high-tension winding very near to metal clamps. Experience had shown that it
was desirable to preserve long-creepage distances and to keep coils of different
phases as far away from each other as possible ; any clamping that had to be done
was done in accordance with old-established rules of insulation. The result was
that in many cases the clamping was insufficient

It will be seen that when we are dealing with forces of many hundreds of
pounds per foot run, especially when many coils are grouped together, very
strong clamps are necessary to keep the windings in position. The old plan of
tying coils together with torpedo twine and securing them with wooden blocks is
wholly insufficient. One cannot hope to make a satisfactory construction without
using strong metal clamps. This necessitates insulating the whole of the winding
outside the slots with an insulation which is not only strong enough to with-
stand the whole testing pressure, but is of such a good mechanical nature that it
will not be crushed under the pressure of the clamps.

We may consider here the various methods of clamping the windings of turbo-
generators.

Where the winding is of the barrel type the clamping may be carried out in

the manner shown in Figs. 130 and 135. Two main objects must be kept in view

in designing this clamping. First, the individual coils must be stayed so that they

cannot move relatively to one another, and secondly, the winding as a whole must

be prevented from being attracted to the nearest part of the frame. One

advantage of the barrel winding is that the field produced by some of the conductors

is to a certain extent neutralized by the field produced by conductors lying near, so

that the magnetic field over a great part of the coils is not as great as with other

types of winding. The magnetic field, however, at the point c in Fig. 130, is as

great as with any other type, and, at this point, the coil is difficult to support.

W^ith this winding the conductors belonging to different phases lie next to one

another, and it is very necessary to insulate the coil throughout its whole length

with insulation strong enough to resist full pressure to earth. It is usual to

THE ELECTRIC CIRCUITS 135

impregnate the coils as a whole and to place them in open slots, the coil being
secured by wedges in the top of the slot. Various methods of securing the
parts of the coils lying outside the slots are given in the paper quoted above.

MATERIAL FOR CONDUCTORS.

Copper is almost universally employed for conductors in the armatures of
electrical machines, because of all commercial metals it occupies the least
space* for a given current-carrying capacity. In addition to this advantage, it
has many excellent mechanical qualities. It is easily drawn to wire and strap ;
its ductility enables it to be bent without much fear of breaking; it can,
moreover, be readily welded and soldered, and the action of the air upon it
does not create any deleterious oxide. It is thus an ideal material of which to
make the conductors in dynamo-electric machines. Its price, however, is high,
and it has been suggested from time to time that other metals— notably
aluminium — might be employed.

The use of almniniom. Some firms are now using aluminium wire instead
of copper wire for field coils. Aluminium is always covered with a thin
film of oxide, and this film can, by certain chemical processes, be made sub-
stantial enough to act as an insulator between successive turns of a wire coil
when the voltage per turn is very low. Thus it is possible to use wire without
any cotton covering, the oxide being relied on as insulator between turns, a
thin sheet of paper being used between the layers of wire to prevent short cir-
cuiting. The conductivities of aluminium and copper are in the ratio of 1 to 1*66
at 15" C, so that if an aluminium wire of the same resistance must be used, it
must have a cross-section of 1-66 times that of the copper wire which it replaces.
If both wires are round or both square, the aluminium wire will have a diameter
29% greater. In the case of small wires, where the cotton covering makes the
diameter so much greater than the bare copper as to allow room for the sheet
of insulation, an aluminium wire takes up less room than the copper cotton-
covered wire.

In many cases it is not necessary to keep the size of the aluminium coil strictly
within the space limits of the copper coil. On machines where there is room
it will often pay to use aluminium, t even though the coil is much more bulky.

It is difficult to give any reliable figures of the comparative cost of copper
and aluminium coils, because the labour is in some cases a large item and differs
widely in different factories. For coils of thick wire such as tramway coils,
where the cost of material is great as compared with the cost of labour, the
saving on the use of aluminium amounts to 25 or 30 % . In the case of small
wire coils of, say, No. 28 or 30 s.w.G., the material will cost 80 or 100 % more

•"Electrical Conductivity of Ck)pper," Wolflf and Dellinger, Amer. I.E.E. Proc,, 29,
p. 1881, 1910 ; ** High-oonduotivity Cast Copper," E. Weintraub, Amer. Elec. Chem. Soc.
frans., 18, p. 207, 1910; "Conductivity of Cbpper," Hirobe and Matsumoto, Elektrot. Ztit.^
33, p. 1245, 1912.

tThe patent rights for the United Kingdom and the Colonies of the Spezialfabrik fiir
Aluminium Spulen und Leitungen, G. ro. b. H. Berlin, are held by the Manchester Armature
Repair Co,

136

DYNAMO-ELECTRIC MACHINERY

for single cotton-covered copper wire than for aluminium wire, occupying the
same room and having the same resistance. But the aluminium wire must be
wound layer for layer instead of "mush," as is often done with cotton-covered
wire, and if this is done by hand the labour comes out rather high. For
cylindrical coils which can be wound by machinery the aluminium coil is
certainly cheaper and in many respects better.

In addition to the saving in cost there are several other advantages claimed for
aluminium coils. The absence of the cotton covering makes the heat conductivity
very much greater than for coils of cotton-covered wire. It must, however, be
remembered that enamelled copper wire* is now very widely used. The heat
conductivity of this, allowing for thin sheets of paper between layers, is almost as
high as the heat conductivity of aluminium with similar sheets of paper.

The good heat conductivity of the aluminium coil leads to a rather lower mean
temperature and to a corresponding lower resistance ; so that, one may sometimes
employ a cross-section considerably less than 1 66 times that of the replaced copper
wire without exceeding the prescribed temperature rise. Thus, in the series field
coils of traction motors, it has been found by experience that in using aluminium
wire it is only necessary to increase the section to 1*4 times the section of copper
When square aluminium wire is employed for field coils, the heat con-

wn-e.

ductivity from wire to wire is extremely good. The insulating oxide withstands
vibration very well, and cannot be destroyed by heat, so that field coils of this
kind can be made which give very satisfactory service. For coils of this kind
the paper between layers is replaced by asbestos. As the weight of an aluminium
coil is only one half that of a copper coil, the handling during the process of
manufacture is very much easier, and some saving is made in freight. In the
case of traction motors! the saving in weight is of special importance. Some
particulars of standard railway motor field coils wound in aluminium are
tabulated below :

Weight of field coil in lbs.

Weight saving

Type of Motor.

Maker.

Copper.
65

AluDilnium.

per car, lbs.

GE 800

B. T. H. Co.

29-5

142

GE 52

f >

21-2

206

GE 58

>»

27 0

264

DK 25a

Dick Kerr & Co.

! 40

18-0

176

W 3a

Westinghouse

29-0

280

GE 66a

/ (i) B. T. H. Co.
l(u)

136

1 57

68-01
28-5J

390

B 17/30

Siemens & Halake

117

1 530

266

There are not many cases in which, at present prices, it would pay to use
aluminium for armature conductors, but if the price of aluminium falls very much
below the price of copper the loss of space may, to a certain extent, be counter-
balanced.

♦"Black Enamelled Wire," EUct, Rev., N.Y., v. 51, p. 611, 1907.

t** Aluminium Windings for Field Magnets of Traction Motors," A. Manage, Lumitre
Electr., 14, p. 104, 1911.

THE ELECTRIC CIRCUITS 137

There are some machines in which the room taken up by the active conductors
is not a vital factor in determining the size of the frame. For instance, iu the
armatures of A.C. turbo- generators, which are external to the field-magnet, there is
usually plenty of room for the conductors, particularly if the voltage is low (see
page 274). A case of this kind is oonaidered below.

Suppose we had to build a 37oO-k.v.a, 25-cycle generator, running at
1500 H.P,M., to deliver 3200 amps, at 650 volts, three phase. We should find that
the size of the machine would be determined by the size of the rotating field-
magnet, and whether we use a small slot or a somewhat larger slot hardly affects
the cost of the frame. Thus there would be no difficulty in finding room for
an aluminium conductor. Moreover, the voltage being low, the cost of the
insulation would be small as compared with the cost of the metal.

As it would not be convenient to provide more than two paths in parallel on a
two-pole armature, each conductor might be designed to carry 1600 amps., and it

no. l&T.— Stnnded alanilniuni conductor to

carry the ume cuitBot m In Fig. 156 trtOi the
ume tempersture rln.

would be desirable to use a stranded conductor, which would occupy the space
shown in Fig. 156. In this size of conductor, we have allowed 1 watt per sq. in.
cooling surface. Now, if a stranded aluminium conductor of the size shown in
Fig. 157 were used, the cooling surface would be increased 28%, and thus would
permit of the use of a conductor having only 44 % greater cross-section than the
copper conductor for the same allowance of cooling surface per watt. It would be
seen that even at the present prices of the metals (copper at £80 a ton and
aluminium at X90 a ton) there is a theoretical advantage in using the aluminium
in this case, and if the prices were reveraed, there is little doubb that the
difficulties at present in the way of using aluminium would be overcome in such
cases as this.

Again, on the rotors of induction motors there is often plenty of room to use
aluminium in8t«ad of copper, and the ease of casting the cage in position warrants
the change. It must be remembered that the mechanical qualities of aluminium are
not BO good as those of copper, and there is not with it the same ea^e in making
thoroughly satisfactory electrical joints. It must be remembered, too, that the
amount of insulation taken to envelop an aluminium conductor is greater than
when copper is employed.

138 DYNAMO-ELECTRIC MACHINERY

High-resistance metals. It is aometimes an advantage to increase the
resistance of conductors without reducing their size, as, for instance, in the
rotors of crane motors of the squirrel-cage type, where resistance is necessary
in order to give the motor the right characteristics, and where considerable
substance is required in the conductors in order that they may not heat up at
too great a rate. In such cases brass conductors have been
employed ; in others, copper conductors are placed in the
slots, and these are connected to rings of high -resistance
alloy.

Shape of conductors. Kound copper wire is most generally
useful where the size is small, and where it is uecessary to
follow tortuous paths necessitating bending in various direc-
tions. The round insulated wire presents no sharp corners by
which to injure its insulation, whereas square wire is only
satisfactory where such control can be kept over the position of the comers
as will avoid danger to the insulation. Where this can be done, square or
rectangular wires offer advantages in giving a larger space factor, better heating

^^9^F?7ttj^

„

•

'pf^^'P^'PT^

TK. 103. FlO. lU.

Method! of arransiiie conduction la slotii.

conductivity, and it is often possible with rectangular wires to make a bett«r
arrangement of the armature conductors so as to keep down the voltage between
adjacent turns.

Arrangement of condnctors in armature slots. In continuoue-current armatures
and other low-voltage armatures in which a " barrel " winding is employed, a.

THE ELEC3TRIC CIRCUITS

139

wSnkn

mtoA

«*.

with compound*

common arrangement of conductors is that shown in Fig. 158. In this case
the conductors near the bottom of the slot form part of one coil, and those at
the top of the slot part of another coil, the general scheme of winding being
as shown in Fig. 505. Where many small conduotors are required, the arrange-
ment shown in Fig. 159 is commonly employed, the coils being wound with small
rectangular copper straps* For smaller wires still, it is better to use round
wires. In many cases complete coils are built up of sections wound on formers,
as described on page 151. In other cases the slot is filled with wires which have
no definite arrangement, as shown in Fig. 44 . This type of winding is commonly
known as a " mush " winding. The copper
space factors for these different tjrpes of
windings depend, of course, upon the thick-
ness of insulation used.

In alternate-current generators and in
induction motors operated at a high voltage,
the arrangement of conductors is sometimes
carried out in the manner indicated in
Figs. 160 to 165. The arrangement shown
in Figs. 160, 161 and 165 is to be preferred,
because in this case the voltage between
adjacent conductors is never more than
the voltage of one turn. The heat con-
ductivity with this arrangement is also
good. In Fig. 161 the two conductors in
each pair, 1, 1 ; 2, 2 ; etc. are in parallel.
Two conductors are used in preference to
one wide one in order to get greater ease
when bending on edge.

The shape of conductors in armatures
is effected by considerations as to the eddy currents which will be produced in
them. This matter is considered fully on page 144.

Arrangement of conductors on fleld-magnets. Some field- magnets are
wound with distributed windings, resembling ordinary armature windings. In
these cases the conductors can be arranged as illustrated in Figs. 133, 215, and 220.
WTiere the field flux of the magnet remains fairly constant, one is not afraid of
eddy currents, and, therefore, the radial depth of the conductors may be made
very much greater than would be permissible in an armature caiTying an alter-
nating current. Fig. 371 shows the arrangement of conductors in the slots of
a turbo field-magnet, in which very deep conductors are used.

In the exciting coils of field-magnets rectangular straps will be used where
the current is very large (100 amps, or more). For smaller currents rectangular
or square wires will in general be preferred to round wire down to size about
•072 in diameter. Below this size it is more convenient to use round wire.
In any case, round wire is often used up to large sizes on account of the
ease with which it is wound, and as the layers of large round wire bed very
well into one another, the space factor is fairly good.

FIG. 165.— Section through slot of an 11,000-volt
S-phaee star-connected generator. Showing best
method of grouping and Insulating for high
stresses.

140

DYNAMO-ELECTRIC MACHINERY

Space fieu^tor in wire-wound coils. The space factor* in wire-wound coils
depends on the thickness of the insulation and the closeness with which the
successive layers are bedded into one another.

In Fig. 166 the space factors of various sizes of round wire with various types
of insulation are given.

Curve A gives the factor for round wire double cotton-covered where the
wires are arranged in square order, and curve B gives the factor where they
are arranged in close order. The curve C shows the possible space factor in

O -05 O-i 0-f5 0'2 ZS O-^ -3J

Fio. 166. — Space factors of different sizes of round and square cotton-covered wire.

coils wound with single cotton-covered wire. Where the coil is not wound turn
for turn, but in mush fashion, the copper factor is much lower, as shown in
curve E for double cotton-covered wire and curve F for single cotton-covered
wire. These curves have been plotted from values obtained from ordinary
cotton coverings. Manufacturers of covered wire supply specially fine cotton
coverings which give rather better space factors than those given in Fig. 166.
Curve D shows the space factor, which can be obtained by carefully winding
double cotton-covered square wire.

* See *' Factors Governing Space Utilization of Electromagnet Windings," C. R. Underhill,
Eltc, Worid, 53, p. 155, 1909.

THE ELECTRIC aRCUITS 141

Where square wires are employed, a better space factor can be obtained,
as will be seen from Fig. 163.

Flat straps have come much into use instead of square wires, because they can
in general be adjusted more nearly to fit given sizes of slots than square wires,
and, moreover, the wider flat faces compel the straps to lie more closely than is found
to be the case with square wires which have received a small accidental twist.

The arrangement of a number of flat copper straps as illustrated in Fig. 1 63 has
-decided advantages over an arrangement of the same number of square wires.
In Fig. 163 no two adjacent conductors have a greater voltage between them than
the voltage of one turn.

SIZE OF OONDUCrrOIlS.

The current density which can be used in any conductor will depend upon
the cooling conditions and upon the permissible temperature rise. The matters
which effect the cooling conditions are — (1) The thickness of the insulation;
(2) the number of conductors assembled together ;. (3) the temperature of the
surrounding iron or air ; and (4) the possibility of heat being conducted away
along the conductor. These matters are considered at greater length in the
<)hapter on "Heat Paths." The designer knows approximately the current
density which can be employed in the type of winding he is employing; and
having taken a conductor of suitable size, he finds the total cooling surface and
the number of watts lost in the conductor,, and adjusts the size until the cooling
•conditions are sufficiently good (see page 222). He knows that, in low-voltage
bar-wound armatures, the current density for 40° C. rise may range between
2500 and 3500 amps, per square inch. In high-voltage armatures the current
density will range from 1500 to 2500, while in shunt coils the current density
may be 1000, 500, or even fewer, amps, per square inch.

The following methods of calculating the size of conductor will usually be
sufficient, though in cases where it is important to cut down the copper to the
smallest possible limit more refined methods of calculation, such as described in
-Chapter X., will be employed.

Shunt coils. The size of wire for the shunt coils of a field-magnet is deter-
mined by the length of the mean turn of the coil and the voltage to be applied
to the coil. The resistance of one turn of the coil must be such that, if the
voltage on the coil were applied to that one turn, the current which would flow
is equal to the total ampere-turns required to be carried by the coil. Any
multiplication of the turns multiplies the resistance and divides the amperes by
the same factor, and as the number of turns is increased the watts required to
give the required ampere-turns become less. The size of wire then is fixed by
the ampere-turns required, the voltage in the coil and the length of mean turn,
while the number of turns determines the watt loss.

The following formulae give the cross-section of the wire required for a given
number of ampere-turns, A.T., a given mean length of turn, /(, and a given
voltage per coil, V.

A.T. X 7-5 X 10-7x12 xT .• r ■ •
= « = cross-section of wire in sq. ms.

142 DYNAMO-ELECTRIC MACHINERY

Or, in centimetre measure,

A.T. xl.7xlO-«xl-2x/t .. r • .
p = cross-section of wire in sq. cms.

The factor 1*2 is introduced on the assumption that the temperature rise of
the coil will be 50* C. above the atmosphere, which is taken at 15' C.

The formulae are also correct if A.T. represents the total ampere-turns on the
poles, and V is the voltage across all shunt coils connected in series.

In the practical calculation of shunt coils the main consideration is the pro-
vision of sufficient cooling surface to keep the coil cool. In Chapter X. some
data are given relating to the cooling of wire-wound coils. For the present
purpose it is sufficient to assume that we know from experience the number of
square inches of cooling surface to be allowed for every watt lost. For instance,
the usual figure with moderate speed continuous-current generators, having only
natural ventilation, is 2*5 sq. ins. per watt, where the permissible temperature
rise is 40** C.

Now we know, from previous measurements on the frame with which we
are dealing, the approximate cooling surface on each shunt coil. Divide this-
surface by 2*5 (if that is the number of sq. ins. per watt to be allowed). We
now arrive at the total watt loss permissible in that coil. Knowing the voltage
at the terminals of the coil (due allowance being made for volts absorbed on
the rheostat), we divide the watts by the voltage and obtain the current. The
number of ampere-turns on the coil divided by the current gives us the number
of turns per coil, and multiplying the number of turn^ by the length of mean
turn we get the total length of wire. The voltage divided by the current give*
us the required resistance of the coil, so the resistance divided by the number
of thousands of feet gives us the resistance per thousand feet. Having obtained
the size of wire, we must see whether the number of turns of that wire can be
put nnto the available winding space. If they cannot, and if the cooling surface
which we have taken cannot be increased, or the cooling conditions improved,
then it is not possible to obtain the number of ampere-turns on the pole in
question without having a higher temperature rise (see p. 504).

Example 21. A certain 250- volt, continuous- current, 6-pole generator, running at

800 R.r.u., requires 6200 ampere-turns per pole at full load. The length of mean turn is

44 inches, or 3*66 feet, and the total cooling surface available is 770 sq. ins. per ooiL

What is the size of wire and number of turns required on the assumption that we are

allowing 2*5 sq. ins. per watt, and a margin of lo % in the rheostat at full load ?

770

—^=308 watts per coil.

The voltage on the whole shunt winding (deducting 15 % for the rheostat) is 212.* Dividing
this by 6, we get 35*3 volts per coil.

3-.g = 8-7 amps.,

6200 -,„ ^ .,

-^-.- = 712 turns per coil,

712x3-66' =2610 feet per coil,

„ « =4*06 ohms,
8w

4*06

-,-.i = l*55 ohms per 1000 feet.

2bl '^

THE ELECTRIC CTRCUITS 143

We now look in the wire table for a wire having a resistance at 70° C, of about 1*55 ohms
per 1000 feet. Now, No. 13 s.w.o. wire 0-92* diameter has a resistance of 1*46 ohms at 70° C.
We might choose this wire and allow for a greater margin on the rheostat, or we may
wind 140 turns of No. 14 wire and the remainder with No. 13 wire, if the cost of making
the joint is less than the difference in the cost of the wire.

We now try if the above number of turns will go in the winding space, which in this
case may be S\"a 1". This would allow 80 turns per layer and 9 layers =720 turns in all.
We now calculate the length of mean turn more exactly, and the cooling surface, and see
that the allowance per sq. in. is sufficient.

Series coils. The size of copper strap to be used in a series coil is some-
times settled from the circumstance that the whole winding must not have
more than a certain resistance. In other cases it is sufficient .that the heat
generated by the passage of the current shall not cause an excessive temperature
rise. In the latter cases a rough estimate must be made of the cooling surface
available, and the size of strap fixed, so that the watts per sq. in. are not too
great. The rules for calculating the watts per sq. in. are given in Chapter X.
(see pp. 230 and 489).

Calculation of the length of mean turn. The only accurate method of finding
the mean length of turn of an armature or field coil of an electrical machine is
from a lay-out on a drawing board, but for the purpose of making quotations on
machines for which no drawings have been made, it is well to have quick simple
rules for estimating lengths.

For barrel-wound armatures having approximately a full-pitch winding, a simple
rule is to add the length of the iron to 1 '4 times the throw of the coil in inches,
and to this add 3 inches. Multiply this by the total number of conductors.

Calcnlation of resistance and weight. An easy rule for getting the resistance
of 1000 feet of any size of conductor (at 15* C.) is to divide 0'0082 by the
cross-section in square inches.

0*0082 ohm is, of course, the resistance of a conductor 1 sq. in. in section
and 1000 feet long.

To get the weight in lbs. per thousand feet of a conductor, remember that a
conductor of 1 sq. in. section and 1000 feet long weighs 3800 lbs., so merely
multiply 3800 by the cross-section in sq. inches.

If we are working in metric units, divide 0*17 by the cross-section in square
centimetres, and we get the resistance of 1000 metres of the conductor. To get
the weight in kilograms of 1000 metres, multiply the cross-section in sq. cms.
by 890.

Example 22. An 8-pole c.c. armature has 76d conpluctors each of 0*06^ x 0*5'' copper
strap. The length of the armature is lOI'', and the throw of the coils 13^ '- Find the
weight of copper and the resistance of the '8-pole lap-wound armature.

ia|+(i3xr4)-i-3=3r7,

31 7 X 768 X tV =2020 feet,
0-06x0-5=0-03 sq. in. "J^ =0*273 ohm per 1000 feet,

2-020 X 0*273 =0*55 ohm all in series,

0*55

-^^ = 0*0086 ohm with 8 paths in parallel,

0-03 X 3800x2*02 =230 lbs. weight of copper.

144 DYNAMO-ELECTRIC MACHINERY

In some cases, as, for instance, in induction motors and generators of high
voltage, the straight part of the coil will project several inches from the slot,
and in these cases we must add more than the 3 inches. In the case of some con-
tinuous-current machines, where there is a considerable length of conductor from
the end of the armature coil to the commutator, something more must be added.
For concentric coils of the type shown in Fig. 112 a simple rule is to add the
length of iron to the pitch of the poles, and to this add A inches, where
A is 12" for voltages up to 1000, 16" for voltages up 3000, 20" for voltages
up to 6000 and 22" for voltages up to 11,000. Where the type of winding
differs from Fig. 114, it is easy to concoct a simple rule of this kind for the
mean length of turn which will give the weight of copper and the resistance
sufficiently near for the purpose of estimating.

The mean length of turn on a shunt coil depends upon the depth of coil
and the amount of space allowed between the pole and the inside of the coil.
In practice, the designer knows from the frame he is using, and the methods
of mounting the coil, how much to add to the two sides of the pole in order to
get the half mean length. Where the coil is a fairly tight fit on a rectangular
pole, it is sufficient to add twice the depth of the winding to the sum of the
two sides of the rectangular pole to get the half mean length of turn.

Example 23. A rectangular pole measures 8{-" x 10 j", the depth of winding is 1 J*.
There are 800 turns of No. 14 8.w.g. wire (0005 sq. in.) and 8 poles on the machine. Find
the resistance and weight of the 8 coils.

8-2o+10-5 + 3=21-75 for the half mean turn,

21-75 X 2 X 800 X 8 X yV=23,000 feet,

nVu^K- = l'W ohms per 1000 ft.,

23x1 -64 = 37 -8 ohms,
0-005x3800x23 = 440 lbs. of wire.

Eddy cuirents in annatnre condiictorB. A very important matter to be
considered when fixing upon the size and shape of armature conductors is the
eddy current, which is induced in the body of the conductor by the magnetic
field set up either by the current in the conductor itself or by the currents
in the neighbouring conductors. We will confine our attention to conductors
placed in armature slots, because the surface-wound armature is seldom em-
ployed. When the conductor is in a slot, the eddy current may be produced
either by a magnetic field which travels down the slot parallel to the length
of the tooth or by a field which crosses the slot from side to side. Except
in those cases where the iron of the tooth is very highly saturated, the magnetic
field passing down the slot is so weak that it does not, in practice, cause any
trouble from eddy currents. In cases where the teeth are highly saturated, the
width of any individual conductor measured across the slot should be kept as
small as possible. The eddy-current loss in watts per cubic centimetre in it can
be calculated by the following formula:

;r, = T.^ X 1 X Q^ 7i2 X bLx X 10-^«,
u p

THE ELECTRIC CIRCUITS 145

where /« is the thickness of the copper strap measured at right angles to B,
n is the frequency and Bmax the maximum flux-density threading through the
conductor. Where the product of ntc is great the induced eddy current interferes
with the impressed B, so that to get a correct result it is necessary to allow for
the change in B due to the eddy, but for values of lUe less than 25 we may
take Bmax at the impressed value. We ma}^ take p for warm copper at 2 x 10"^\

It will be seen from the following example that this loss is usually of very
little importance :

Example "24. B,n»»=400 (corresponding to about B =20000 in the teeth),

rt=50,
/«=-25,

>r*=8-2 X 0-0625 X 2500 X 160000 X 10-".
=0*00205 watt per cii. cm.

Where, however, the teeth are highly saturated (say to 25000), and the
conductors are thick, the eddy-current losses at 50 cycles become appreciable.
This matter has been investigated experimentally by Dr. Ottenstein, and for
further information the reader is referred to his paper,^ which deals also with
losses due to flux fringing from the sides of the teeth.

The eddy current which is produced by the magnetic field which passes
across the slot, and which is usually produced by the current carried by the
conductors in the same slot, may be very great indeed if the conductors are
not properly designed. This matter has been fully dealt with in a paper t by
A. B. Field, read before the American Institution of Electrical Engineers, in
which the theory is fully worked out, and some very useful curves given by
means of which the eddy current in any case can be arrived at in a very simple
manner. Figs. 167 and 167a are reproduced by permission from Mr. Field's
paper, and examples showing how they are employed are worked out below.

The amount of the eddy-current loss is a function of the radial depth of
the conductor, and also of the current which passes in the slot between the
point at which the eddy current is being considered and the bottom of the slot.
When there are a number of conductors one above the other, the conductor
nearest the mouth of the slot, being in the strongest field, has the greatest eddy-
current loss. In Fig. 167 the curve marked m^ refers to a conductor nearest
the bottom of the slot; m^ refers to the conductor next to it, and so on,
the higher numbers of m being nearer the mouth of the slot. If there is
only one conductor, the curve m^ refers to it. The eddy-current loss is also

* ** Das Nutenfeld in Zahnarmaturen iind die Wirbelstroniverluste in massive Armatiir-
Kupferleiteni," Sammlung electrotechniecher Vortrdgt^ Stuttgart, 1903.

\Proc, Amer. Inst, Elec. Engrs., vol. 24, p. 761, 1906. See also Electrical World, vol. 48,
29 Sept., 1906, where some experiments are described which corroborate the conclusionH
arrived at by theory. In the Jmimal of tht InstUtUion of Electrical Engineers, vol. 33, p. 1125,
the matter is still further elaborated by Mr. M. B. Field, and some practical cases considered.
** Eddy -current Losses in Armature Conductors," Oirault, Ltimiere Meet., 4, 35, 1908;
" One sided Distribution of Alternating Current in Slots," Kmde, Elek, uwi Maschinenbau,
26, pp. 703, 726, 1908; **Skin EtesisUnce Losses in Alternator Windings," F. Rusoh,
EUctroteck, u. Maschinenbau, 28, pp. 73 and 98, 1910 ; " £ddy -currents in Solid Armature
Conductors," Angermann, Mekt. u. MaschiiienlKnu, 28, p. 975, 1910; '* Copper Losses in a.c.
Machines," RogoM-ski, Archiv f Elektrot,, 2, 81, 1913.

w. M. K

146

DYNAMO-ELECTRIC MACHINERY

H function of the ratio of the width of the copper to the width of the slot.
This ratio Mr. Field denotes by r^. If a conductor consists of a number of
parallel straps one above the other, which are soldered together only at their

2.7

VALUES OF af
2.6 2.5 2.4 2.3 2.8 2.1 2.0 1.9 1.8

1.7 1.6

1.5

1.4

S.0
8.9
8.8
8.7
8.6
8.5

8.4

8.3

8.8

U.8.1
O

< « A

1.8
1.7
1.6
1.5

Jf\^

f \

1.4
1.8
IJ
1.1

^\y

)

I .;

S .4

s .(

r A

8 .1

% 1

.1 1

VALUES OF af

FIG. 167. — OlTTing ratio K^ of tota] loss with eddy current to normal loss without eddy current
in armature conductors. For use of fractional values of m see pages 110 and 150.

outer ends, then the eddy-current loss is also a function of the ratio of the length
of conductor lying in the slot to the total length between the soldered ends.
This ratio is denoted by r.,. The eddy-current is also a function of the frequency n.

THE ELECTRIC CIRCUITS

147

We will denote the quantity 0'145Vwr^-f r^ by a and the depth of the con-
ductor in centimetres by /. Then the Values of the product af can easily be
found for any case. The ordinates, Kdj of the curve give the ratio of the actual

2.7 2.6 ftJ& 2.4

VALUES OF af
2.8 2i2 2.1 2.0 1.9

1.6

l.r 1.6 1.5 1.4

1—

(

■~-

ni
^

/ / /

u.
O

IU6

/ /

>
6

/ \ f

00/
Ml *i

^ /

»/

V 1

7* ^.

/ 1

■y.

li)

•

9 .]

• 4

t Z A .1

\ .d

\ .l

.«

\ .s

» 1.

VALUES OF af

no. 107a.— Qiving ratio Ka of total loaa with eddy cumnt to normal loss withoat eddy cnnent

in armatore conductors.

copper loss to the copper loss there would be if there were no eddy current. The
abscissae of the curves are the values of af. We give below a few illustrations
showing how the curves are employed.

148

DYNAMO-ELECTRIC MACHINERY

Example 25. Suppose that we have a slot '5" wide, containing a single conductor '236"
wide and 1" deep. As the conductor is solid, rg equals I and r^ equals *47. Suppose that the
frequency is 50, then the value of a is '706 and /=2-54. Then the value of 0/"= 1'79. We see
from the curve that for a value of a/*=l"79, A'd=l*67. That is to say, that with a conductor
of this depth working at 50 cycles, owing to the eddy current, the loss is 67 % higher than it
would be with a continuous current passing through the bar.

For a frequency of 25 cycles, the value of of works out at 1*27, from which we find
that Arrf = l-23.

Example 26. Fig. 227 shows the conductors of a 25-cycIe generator drawn full size.

'45
Here ^i = roiQe= *553. Although each conductor is divided into two parts, these parts are

sweated together at no great distance from the end of the slot, so that we may take r^ &a

practically unity. We therefore get a = •145\/25x -553= '54. And a8/= 1 '27, we get 0/= 0'686.
Referring now to Fig. 167, we find K4

for m4^1-88
for wig =1*47
for wi,= l'17
for mi = l'02

4 JS^

Therefore the mean value of K4 is 1 *38. That is to say, that on the generator in question
the loss in the conductors lying in the slots is 38 % greater than it would be if there were no
eddy currents.

Example 27. Suppose that we are designing the armature of a 50-cycle turbo-generator,
in which it is required to have two conductors per slot, each to carry 700 amps., and that
the maximum width of slot permissible is *025", and the room required for finish and insulation
is '15", so that the copper bar cannot be made more than |" wide. The question arises aa
to what is the best depth of bar. If we were not concerned with the eddy current, we
might begin by assuming a current density of about 1900 amps, per square inch, which
would require a bar about |" x §*. Let us firat arrange two bars |* x §" with the required space
for insulation. It is easy to show from Mr. Field's curves that this arrangement of con-
ductors not only requires more copper than is necessary, but the eddy currents make
the total losses in the conductors 35% more than they would be if the conductors were
reduced to the best size. In order to find the best size, it is a good plan to plot the
figures proportional to the losses in each conductor for different depths of copper. For
this purpose it is convenient to work in centimetres. We have here 0= •145v50 x '675= '845.
It is convenient to write down the calculated values in columns shown below :

/..

a^fl.

A'd.

'A '

A-

CJ\.

A'rf.

3-4

1'6

1-35

1'27

•795

1-6

1-35

1-8

1-52

1-395

•773

1'4

118

2'39

2-0

1-69

1-56

•78

1-2

101

1'78

2-2

1-86

1'75

•795

•845

1-39

2-4

2-025

1-92

•8

•675

116

2-6

2 '2

215

•825

■6

•505

1'05

K4
ft'

2
1
1
1
1
1

485

The reduction on the depth of the top conductor will give more room for copper in
the bottom conductor, and therefore, as /,, the depth in centimetres of the top conductor
is reduced as shown by the figures 1'6, 1*4, 1*2, etc., the depth of the bottom conductor
is increased as shown by the figures 1'6, 1*8, 2*0, etc. The resistance, apart from eddy
currents, is inversely proportional to the depth of the conductor, and so we have divided
K by /i in order to get a figure proportional to the losses. This is done in the 4th and
8th columns. Plotting the values of these columns, as shown in Fig. 168, we see that the

THE ELECTRIC CIRCUITS

149

losses in the bottom conductor are almost constant, while the depth is changed from 1*6
to 2*6 cms., the minimum occurring at 1*8 cms. The losses in the top conductor reach a
minimum at about 1 cm. depth. It must not, however, be supposed that this depth is
the best, because the cooling conditions on a shallow bar are not as good as on a deeper
liar. Moreover, there is always some length of bar outside the slot, in which the eddy-
current loss is not so important, and in this part it may be desirable to have a deeper
Ijar (see Fig. 438). In those types of windings in which it is convenient to use a connector
larger than the bar in the slot, the value chosen for the depth of the top conductor would be
rtither greater than 1 cm., say '4". This would give a current density of about 2300 amps,
per square inch. The value chosen for the depth of the bottom conductor would be about
rS cms., say 'T. This would give a current density of 1620 amps, per square inch. The
losses per foot nin in the two conductors together, when hot, amount to 45 watts, and
as the cooling surface per foot amounts to
nearly 45 square inches, we have an allow-
ance of 1 sq. in. per watt, which is quite
sufficient for mica and paper insulation
more than ji" in thickness (see page 222).

The alternative arrangement would be
to use a stranded conductor in the top
slot and a solid conductor in the bottom.
The stranded conductor could then be
made of greater cross-section, say *7 x *625,
having a total section of *38 copper. This
arrangement would give a lower tempera-
ture rise on account of the reduced eddy-
current loss in the stranded conductor, a
larger section of copper permissible and
the greater cooling surface. A stranded
conductor, however, is not as strong
mechanically as a solid conductor, and
is rather more expensive.

ExAMPLB 28. What is size of copper
Htrap to put in the armature slots of a
3-phase 50-cycle generator running at
1000 ILP.M., if the voltage is 500 and the
current 400 amps, per phase ? There are
to be 144 slots having a pitch of 'T and
1 strap per slot. Under these conditions
it is not good Ui use a strap much wider
than j|th inch, which when insulated will

Depth of CorvUuctor m^ in cms.

Fig. 168.— Curves Bhowing how the losBes in oonduclors
change as the depth of the oondncton is changed.

require a slot '28 wide. From previous experience we know that a single conductor of
this kind can be worked at about 3500 amps, per square inch. Let us take provisionally
a strap i" x V. This will have a resistance (hot) of 8x 10~^ ohm, giving a loss of 13 watts
per foot run. Allowing about 50% extra loss for eddy currents (see Ex. 25), we have
roughly 20 watts. Now the area of the surface per foot run is 27 inches, giving us
1*35 sq. in. per watt. This is more than we would require if the iron is reasonably cool.
Suppose that we are allowed 50° C. rise in the copper, and that the iron of tlie teeth rises
only 30** C, then 1*0 sq. in. per watt would be sufficient allowance for cooling. If the
efficiency' guarantees will permit it, we can reduce the bar to ^"xfy thus somewhat i*educing
the eddy-current losses. The resistance will now be 1 "07 x 10"* ohm per foot, giving a
loss of 17 watts, and adding 25 % for eddy -current losses (see Ex. 25), we have 21 '5 watts.
The surface of the foot of strap is now 21*3 sq. in., which would be just about sufficient.
Note that the strap is now worked as high as 4250 amperes per sq. in. After having arrived
at the approximate size by this rough calculation, a more careful calculation may be made of
t he eddy-current losses. In this case a = '682, so for a depth of 2*9 cms. , A'^ = 1 '24. The size of
the end connectors would next claim attention. The cooling conditions here depend greatly
on the fanning action of the field-magnet, on the amount of insulation on the connectors,

150

DYNAMO-ELECTRIC MACHINERY

1 the anioiint of spaoo between thetn. In any case it is DOt worth while to work tlie
I up to the maximum allowable l«inperature. A oonncotor J'x 1" would have a
loas of 13 wattB per Coot run, and us the eddy-current loasea would be small there would
be about 2 aq. in. per watt. This would be suffioient.

If the machine had 72 Blots and were barrel wound (see Fig. 129} with two conductors
per alot, the eddy-ourrent Insa in the conductor in the top of the slot would have to be
seriongly considered. Tho slots could now be about 0-4r wide, allowing room for a bar
0'2S' wide. Taking again the provisional value 3500 amperes per aq. in., we would try a
bar i' deep, and then reducing the bar in the tap slot and increasing the bar in the bottom of
the slot in Bucoeaaive stages we would find the best croas-section by plotting curves as shown
in Fig. 168.

Laminated conductors. Where the conductors are laminated, the laminae
lying in the direction of the Rax crossing the slot, the eddy current may be
reduced. In most laminated windings there are points
at which it is necessary to sweat all the laminae together
for the purpose of making joints, and the amount of
eddy-current loss will depend upon the position of the
sweated points with respect to the conductor lying in
the slot. For a bar winding with laminated conductors
the value of K^ given in the curves applies for the
whole length of the conductor between points at which
the laminae are connected together. The value r, (the
ratio between the length of conductor lying in the alot
to the length between sweated points) introduced into
the formula given above makes the necessary correction
to allow for the lamination. For a. winding in which
the laminae are continued front layer to la,ycr and onlj'
joined together at the beginning and end of the coil,
we obtain Ka applicable to the whole coil by taking
ni = 0'5 + hal[ the number of layers per slot, if the winding is one in which there
are twice as many slots as coils. For the case in which each slot carries parts of
two coils, one above the other, we take instead the curve for which m = 0-5 + one
quarter the number of layers per slot. For a one-layer winding, in which the
conductor is twisted over in the middle of the coil so as to reverse the order of
the laminae, we take 7» = 1, but refer to a point corresponding to 0'5a/ instead
of a/. The reader should refer to Mr. Field's paper for full information upon these
matters.

Stranded condnctors. The main objection to using stranded conductors is
that they are mechanically weak. In eases where the conductors would have
heavy eddy currents in them, if solid they should be made with a twisted strand
and proper means provided for supporting them. Where two conductors greater
than ^ inch are used onu above another in a slot on a 50-cycle machine, the losses
in the conductor nearest the mouth can be greatly reduced by stranding it as
shown in Fig. I68o. The solid conductor can be used to give support to tho
stranded one.

CHAPTER VII.

THE DESIGN OF ARMATUKE COILS AND THE FORMERS UPON WHICH
THEY ARE WOUND.

As we have seen on page 89, armature coils may be broadly divided into two
classes — (1) Concentric eoils; (2) lattice coils. On continuouH-current armatures
and on all machines in which It is important to preserve a uniform step in phase
between one coil and the next, the lattice or overlapping coils are employed.
These overlapping coils are of various types.

Armature coils of copper wire may be of the diamond shape shown in
Figs. 131, 169, and 170, or of the short type shown in Figs. 133 and 171, or

it >t tbe bottom.

the involute type shown in Fig. 136. The involute type, however, is now rarely
used in rotating armatures.

In Fig. 173, on the leftrhand aide, we have a aingle-turn coil of copper strap
arranged for two coils per slot. For a multiple-wound c.c. armature this will
form part of a lap winding.* On the right-hand side we have a two-turn coil
of copper strap, and in Fig. 192 is a strap with the requisite bends in it, made
in a bending machine before the coil is formed.

Coila of the short type (Fig. 171) are usually wound in a former or mould.
Coils of the diamond type may be either made to the correct shape at once by

strap ooil for a ieriee-wound c.c. armature. This will form

152

DYNAMO-ELECTRIC MACHINERY

being wound on a foi*mer, or they may he wound in the form of a simple

loop, as shown in Fig. 169, and "pulled*' to the required shape in a pulling

machine (Fig. 177).

Before proceeding with the actual
design of the formers, on which the
various types of coils are wound, a
few remarks are necessary on the
former or mould itself and various
terms used in the design. The
mould, when required for coils of
copper wire or ribbon or for field
coils of strap wound on the flat, is
made of some hard, well-seasoned
wood, lined on those parts where the
coil is wound with fibre. Fibre is
used because it is easily machined to
the desired shape, and while it wears

well on the mould, it docs not injure the cotton or silk covering of the wires.

The mould is usuall}' made in two parts, to facilitate the removal of the coil,

and sometimes it is necessary to have

Fio. 171. — Short-type armature coils showing how three
" single " sections are grouped to form a " complete " ooU,
also showing method of insulating by interleaTlng the
sections with black mica-doth and wrapping it around the
complete coll.

Fig. 172. — Armature colls of copper strap for lap- winding.

three or four parts which separate in

different planes. Fig. 174 shows a

mould for an armature coil, and the

way in which the two parts of the

mould are separated. One part is

fixed to the face-plate of a lathe

used in the winding and the other

is fixed to the first by means of a

pin and tapered key or cotter. The

pin goes through both halves, and

the cotter, when driven home, holds

them firmly together. The sketch

(Fig. 193) shows how the wire is taken around the various corners.

The mould for a *' pulled" coil is of much simpler design, as can be seen

from Fig. 175.

Where the armature or stator coils
have to be made of stout copper
strap, the mould is usually made of
cast iron. If only a few coils are
required, the mould may be made of
wood with iron fittings. Fig. 176
shows a former designed for making
coils of copper strap like that depicted
in Fig. 173. For this type of former
a winding lathe is not required, the

Fio.173.— Armature coil of copper strap for wave- winding. Straps being simplv CUt to length

THE DESIGN OF ARMATURE COILS 153

and hammered to shape. Any specially difficult bending, such as liending on edge
around a small radius, is done on a bending machine before the strap is put on

FIO. 174. — Uould tor a ■hort-type ■nnstun coll diawing how Ihe two lulvea iie sepanled.

the former. Formers of this kind are usually made adjustable in length, so a
accommodate coils for different lengths of armature.

In winding the coils on the mould, the wire or strap is hammered into position
Uy means of a mallet and fibre drift. The latter is made of various sizes luid

154 DYNAMO-ELECTRIC MACHINERY

shapes, so as to fit into the crevices of the mould and to level down a number
of wires lying together.

In general, the moulds for armature and stator coils are so designed that the
coils, when assembled on the machine, will lit together and make a construction
which can resist the mechanical forces to which they may be subjected, and at
the same time provide for good ventilation.

In case of an armature and stator coil, the slot portion has an extra thick
wrapping of insulation to withstand the voltages to ground, and this straight
portion of the insulation, commonly called the "cell," projects straight out of
the slot a certain distance depending on the voltage of the machine. For machines

CI [or making the type ol coll llluBtrnted In Fls. ITS.

of voltages up to 500 or 600 volts, the projection of the cell is usually about
^ inch for wire coils and about J inch for strap coils, though in very small
machines it is sometimes reduced to ^ inch. The amount that the cell projecte
depends upon a number of considerations. Where the coil is of strap, it is usual
to tape the individual straps well round the corner to avoid the risk of short
circuits at that point, and the straight cell from the slot is made to overlap the
tape ; this generally requires about j inch of projecting cell. Cotton-covered wires
with an insulation of uniform thickness along the entire straight part will not
require more than ^ inch of projecting cell. In small machines wound with cotton
or silk-covered wires, the projection may be cut down to j inch. The amount that
the cell projects can only be cut down with safety in armatures which are
completely impregnated in a vacuum tank (see Table VIII., page 172).

With regard to the actual winding of the coil on the mould, this may either

THE DESIGN OF ARMATURE COILS 156

be done in " sections " or as a complete coil, depending on the shape. For instance,
suppose we have a short-type coil designed to give 30 wires per slot. Half of
theae will be in the upper limb of one coil and the other half in the lower limb of
another coil. That is to say, there will be 15 wires per coil. The bottom half
of a coil will be in the bottom of one slot and the top half of the same coil in
the top of another slot some way further round the machine, depending on the
pole pitch. The winding of each coil, let us say, consist« of 3 x 5 wires, and
tb& mould may be designed so that the coil is wound to shape in 3 "sections,"
each of 5 wires, which, sections are afterwards assembled to form a "complete
coil," Fig. 171 shows 3 sections, which are afterwards assembled to form a

Fio, 177. — Pulling machine for maklne dlunond-ahaped tnnature coUs.

complete coil. On the right-hand side of the figure is seen the completed coil,
and at the back is a coil with the insulation mica and paper tucked around the
inside section and ready to be folded around the 3 sections together. The
"complete" coil may be wound in the first instance with 15 turns, or 5
turns of 3 wires in parallel, or, as desired, the winding being done in a simple
straight mould making a straight coil like that shown in Fig. 169. When the
complete coil is wound in the first instance it cannot be conveniently made in
a shaped mould. It must be formed after winding. This forming after winding
18 generally done on a pulling machine, illustrated in Fig. 177. Where the coils
are wound in sections, the wires all follow on without any crossing in the section
itself ; but if it is necessary to connect the sections in series, there will be a
cross-over from one section to the next if all the sections are wound the same.
Fig. 180 shows a coil with a cross-over. To avoid this, some of the sections

156

DYNAMO-ELECTRIC MACHINERY

Fig. 180.

FIG. 180a.

{one-half where there are an even number of sections and less than half where
there are an odd number of sections) are wound with the mould reversed on
the face of the lathe, so that when the sections are assembled the end of one
section comes directly into line with the beginning of the next. The way of

assembling two coils (one reversed)
to avoid a cross-over is shown in
Fig. 180a.

The "throw" of the coil is con-
veniently expressed by giving the
numbers of the slots in which the
coil lies. Thus we may speak of a
throw of 1 and 16. This gives a coil-
pitch of 15 slots. In cases where
the coil is wound to shape, the mould
is split along the throw-line, i.e. the
line joining the inside top corner of
the bottom half of the coil and the
inside bottom corner of the top half
of the same coil (see Fig. 181). This will be understood more clearly when we
deal with an actual example.

Before laying out any stator, rotor, or armature coil, the following paHiculars
must be known, and should be filled in on the design sheet:

(1) Particulars of insulation, its thickness on the ends of the coil and on
the slot portion, and also the length that the cells must project from the ends
of the slots, the amount that the top cell is to project beyond the bottom one
and the minimum distance allowed for electrical or mechanical reasons from the
ends of the coils to the nearest metal. This is all covered by the insulation
specification.

(2) Bore of stator or diameter of rotor or armature.

(3) Length of iron.

(4) Number of slots.

(5) Size of slots (this is given not as punched size, but as finished size, due
allowance having been made for irregularities in the punchings).

(6) Number of coils.

(7) Windings per coil.

(8) Size of wire.

(9) Throw of coil.

(10) Depth of holding-down wedge or baud-groove.

The designs worked out here cover the most common cases. The other tjrpes
of coils in use can be worked out on the same general principles.

The design of a diamond-type coil for the stator of an induction motor.
M^ith this type of winding there are as many coils as slots, and all the coils
are identical. We shall call that half of the coil that lies in the bottom of the
slot " the bottom half " and the part that lies in the top of the nth slot, further
round on the machine, "the top half." The bottom halves all come straight
out of the slots for a certain distance, depending on the voltage, and then bend

THE DESIGN OP ARMATURE COILS

157

round to make such an angle 6 with the iron that they all fit closely together
and form a surface which is nearly cylindrical (see Fig. 131a). The coils then
bend up and over and form a second cylindrical surface, in forming which they
again fit tightly, and finally they bend to go into the required slots. The design
of the coils is usually the same at each end of the machine. For voltages up

Fig. 181. — Showing the way of laying out the dimensions of the mould for a diamond-shaped coil.

to 500 the coils all carry the same insulation on the ends, but for higher voltages
there will be extra insulation on the coils at the end of each phase, where they
lie adjacent to the coils of the next phase. An allowance should be made for
this in laying out the coils of a high-voltage machine. Ordinary stator coils
of this type require no extra support, but if the projection of the coils beyond
the slots is very long or if the throw is very long, as in two-pole machines^
it is advisable to tie the coils to an insulated metal ring which embraces the
whole winding.

158

DYNAMO-ELECTRIC MACHINERY

The mould worked out below is one on which the coil would be wound to
shape in sections. The particulars of the machine are given on design sheet
No. 1.

The upper part of Fig. 181 shows the shape of the coil as we look at the
stator from the end, and the lower part of the figure as we look at the coil from
the centre of the stator towards the outside ; the lower figure need not be drawn
when laying out the coil.

The coil lies in slots 1 and 13, so the portion of the stator containing these
two slots must be drawn in.

Describe first of all a circle (or part of a circle) representing to scale the

bore of the stator. The angle a enclosed by the centre lines of the slots in

which the coil lies is

1^-1

^x360 = 45^

n-l

or generally — ^- x 360, where X is the total number of slots and 1 and n the

throw of the coil. The cord subtended by this angle at the bore of the stator
is equal to the bore of the stator multiplied by sin Ja. Marking off this chord
on the bore and drawing lines from the centre through its two ends, we get
the centre lines of the slots. The two slots can then be drawn in and also
the circle afc, where the top and bottom portions of the coil touch, due allowance
being made for any packing wedge there may be at the bottom of the slots,
and also for the holding-down wedge at the top.

In the present case no packing wedge is required, as the coils just fit the
slots nicely. The holding-down wedge is taken as coming I" below the top
of the slot. Next draw in the slot portions of the coils as a rectangle (the
various wires need not be shown), allowing the necessary clearance all around
for the insulation specified. Then subdivide this rectangle into as many parts
as there are sections in the complete coil. The end view of only one section need

be drawn, since all the sections
are alike. We will take the one
lying in the right-hand bottom
part of the slot No. 1 and the
right-hand top half of slot
No. 13.

Join the points a^, Og, ue, the

top and bottom inside comers

of the bottom and top halves of

the section under consideration

by the straight line. This is the

" throw line " referred to above.

Turning now to Fig. 182, let t be the thickness of the insulated coil and

// the smallest pitch of the slot on the cylindrical surface formed by the coil

ends, then 0, the angle which the coil makes with the iron, must satisfy the

equation 8in"^= In laying out the mould in the shops it is not very

FlO. 182.

THE DESIGN OF ARMATURE COILS 159

convenient to deal with angles, so, instead of specifying 0, we give the lengths
X, y and Z>, and from Fig. 181 we see that y = B-Xy

or xt&nd^^^D^ytAud^ + l"

^BtaLne^-xttLnS^ + l".

Therefore a;(tan 6^ + tan 6^) = ^ tan ^^ + J''

^tan^^ + j
^^ ^"tane/g + tan^i*

Therefore, if we find the dimensions x and D and lay the mould out accordingly,
the angles ^^ and O^* ^hich the upper and lower coil limbs must make respectively
with the iron, will be obtained.

We now require to know the total thickness of insulation on the ends of
the coils (exclusive of the cotton covering of the wire), and we will assume in
this case that the taping at the ends, including varnish, =0*06" (this is what
it would be for a motor up to 500 volts). This insulation allowance is added
to the thickness of the cotton -covered wires lying side by side.

Now sin ^1 = — and sin 6^ — ~» J^i ^^^ P^ being chosen at points along each
P\ P%

limb of the coil end which is nearest to the centre of the machine, because there

the pitches are smallest, and therefore the angles 6^^ and 0^ the biggest, and the

coils designed to fit at these points will not bind on the parts further from the

centre.

On the calculation sheet the pitch p^ has been taken on the radius R^ and the
pitch |>2 ^^ ^^ radius 72^.

In figuring out the pitch, we have

pitch = -^,

and the rest of the figuring on the calculation sheet is to find x and D.

It has been found advisable in practice, after getting the length x tan ^^ or
^ tan ^1 + -25", to add •125'' to the result, so that

a; tan ^2 + '1 25* = i> = y tan ^1 + -375.

This '125" is added to allow for the cutting back of the end and side bevels,
and increases the angle 0^^ but not 0^.

We can now proceed with Fig. 181, having found that

a; = r and y-Z^-^i^TSS" - 4" = 3-33.

From ag lay off o^i = y along the throw line, and through iVj draw a normal
to it, cutting the bore circle at N^. The part of the coil on the right-hand
side of this normal must never come nearer to the surface of the iron than
•125", and at the same time the coil should drop as little as is practicable below
the iron. It is impossible to wind this part of the coil with a bend in it (as
the winding starts on the outside of the curve and progresses inwards, and
could not therefore be kept in place), but, on the other hand, if it were made
straight, the coil would lie on a tangent and, in many cases, drop further than
is necessary below the iron. In such a case a point iVg is taken on the normal

160 DYNAMO-ELECTRIC MACHINERY

(0*75" maximum below the iron), and through this point a line is drawn touching,
the circle whose radius is E^, If the shape thus given necessitates a bend in
the coil, the bend must be given to it after winding. In the present example
y falls just 0-75 " below the iron.

Make iVgiV^ equal to the width of the uninsulated coil and A^^^V^ equal to-
Y ; this ;J" keeps the two parts of the coil ends apart and enables the coils to
lie consecutively without bending at the turn-over. The other limb of the
coil is then drawn in through N^Nq in arcs of radius M and R^ struck from
the same centre, and the nose of the section can then be drawn with a thickness-
equal to that of the cotton -covered wire.

Oalculation Sheet of Diamond-Sliaped Ooil.

Mould No. 3-241.

Angle between slots = ^^^ x 360 = 45".

The chord on the bore of the 8tator= 18-094 x sin -*# = 6 Q-r, and from Fig. 181, B{ = iher
length of the throw line) = 7'33''.

Thickne8R of coil on ends (including insulation) = 3 x -093 + '06 = -339".

P, lit radius ^1=*^^'=^''^ — =0-601"

and P, „ /?,='2^J:?=?^:Zi=o-636\

s'n^i = -^ = 0-564" and sinda = -|^= -533",
* -601 •* -636

^, = 34 -4" and ^2=32-3°,

^J5tan<?i + 0-25" 7-33 X -685 + 0-25 .^
''^~ tan^a + tan^i "■ -633-1- -685 " '

y = (7-33-40)=3-33,

i)=x-tontf2+125"=(4x-633)-l-125=2'655.

We now have all the information we require for filling in the design sheet
No. 1, which gives the particulars from which the mould is actually made. A
list of these particulars is given below.

Length of cells. The short cell, that is the cell at the bottom of the slot, is.
equal to the length of iron plus twice the projection of the cell beyond the
iron. The long cell, that is the cell at the top of the slot, is equal to the
short cell plus twice the distance that the long cell must project beyond the-
short, according to the insulation specification.
, Wire space — number of turns per section x the size of the insulated wire.
A = length of short cell + twice radius F.

i? = throw of coil. Sometimes, when the straight limb of the coil is bent
after winding, it is advisable to add a little to this, because, after
bending, the throw is a little less than before.
C = ;r.

B is obtained as above.

A' = length of the long cell over the short cell at each end.
jP=the radius put on the corners of the mould to protect the insulation ou
the wire.

THE DESIGN OF ARMATURE COEUS

161

H^ the amount added to ^ at each end, in order that the short cell may be of
the proper length, and is usually about y. If no allowance were made,
this cell could not be made long enough, owing to the intersection of the
side bevel with the end of the mould. This can be seen from Fig. 181.

Bevel N is represented by the rectangular components of a length of the
radii bounding the angle a along the throw line and at right angles to it. The
bevel on the other side of the mould is the same, but cannot always be made
so, owing to the fact that on this side the coil is being wound down the bevel,
And if it were made steeper than 1 in 2, would give trouble in winding. It is,
therefore, usually made with this amount of slope (unless it actually figures
out less), and any extra bevel required can be given by hand to the coil during
the operation of putting the coils in the slots.

The mould can now be made from the dimensions given on the mould sheet.
The proper thickness of wood and fibre to use is found from actual experience.
In the present case the thickness of the mould from back to front would be
about 4" and the fibre lining around which the coil is actually wound Y-

The mould itself is made in two parts, to facilitate the removal of the
section after winding. The split is made along the throw line and parallel to
the axis of the machine.

Design Sheet No. 1.

Specification for Aimatnre Mould.

Motdd No. 3241.
Order No. 78921.
25 H.P. Motor,
2 Phase,

6th May, 1911.
For Electrical 8pec%JuxUion, No, 632.
400 Vdts, 940 R.P,M. 6 Poles.

50 Periods.

Diameter of Armature, 18^ •

Length of Armature, 7? •

Number of Slots, 96.

Size of Slots, §f X 1^*.

No. of Coils, 96.

Winding per Coil, 3 sections each of 5 wires.

Size of Wire, *081" d.c.c.

Coils in Slots, 1 and 13.

Lengths of Cells, 8^ and 9".

Wire Space, |:4*«

^ 8^ • K 1-^ '

c 4r.

T 1"

i>2f
1"

jw^9r.

N 1.2.

Insulation Spec. 1890.

i«

Note. — Depth of holding-down wedge -g

Fig. 188.

W.M.

162

DYNAMO-ELECTRIC MACHINERY

TlQ, 184.

A sketch is made (not necessarily to scale) (see Fig. 184) showing the distances
that the coil projects beyond the end of the iron and how far it falls below the
bore of the iron, so that it may be seen that it does not foul the end bell or

any other part of the machine,
a = length of long cell over iron.
& = length of short cell over iron.
c = drop of coil below bore of the iron.
This can be measured directly off
Fig. 181. A margin (in this case
Y) should be added for safety^
as this type of coil can easily
be distorted.
d = length of short cell over iron + radius F + D + width of coil + length of
stub + a small allowance for safety. The stub caused by the jointing^
of the wires of the different sections will vary with the size and
number of the wires. In some cases it will be possible to bend it
over or to get it between the coils.
In the present case a^Vy 6 = ^, 6 = 2", d^i^".

In the case of a diamond-type coil for a revolving armature, slight modifica-
tions must be made in the procedure. For instance, the bent limb of the coil
will be the one to fall only a short distance below the top of the iron, and the
minimum diameter on which the straight limbs fit together can easily be found
by trial and error. What has

'ZZ3

FIG. 192.

been called the bent limb of the
coil can also be wound straight,
but the finished appearance of
the armature is then not quite so
pleasing. Straight limbs build
up on a curved surface at the
ends. The hollow curve is
sometimes useful for prevent-
ing a band of steel wire from
slipping off.

In the case of strap coils consisting of two or more turns, the mould is
made of iron, and is often adjustable in length by means of a screw (see Fig.
176). On the mould sketch the length of mean turn is given approximately,
and this length of copper strap is taken and bent in a bending machine into
the shape shown in Fig. 192. As a greater length of copper is required for
one side of the coil than for the other, some extra length is allowed, making a
sag at one side, as seen in Fig. 192. The strap, then, is put on the mould, a
pin going through it at each end. It is next bent roughly to shape over the
mould, and the two ends of the mould are screwed apart while the coil is
hammered to shape. Any modification in the length which is then found neces-
sary is made and the other coils formed from a suitable length of strap.

In the case of a one-turn strap coil, the copper is bent into a rough U-shaped
piece and then hammered to shape on the mould.

THE DESIGN OP ARMATURE COILS

163

Where the coil consists of two or more straps side by side, each consisting
of one turn, as is often the case in direct-current machines, the coil is formed
to shape, and while on the former, the straps are opened out at the front end

Fig. 198. — Sketch of a mould for a " short-type " ooil ahowhig ooll in position.

by means of small shaped wedges driven in between, so as to shape them correctly
for lying in the right commutator bars.

The deflign of a short-type coil for a 0.0. armature. We will now con-
sider the design of a type of coil which is sometimes called the "short" type,
because it does not project as far horizontally beyond the iron as the diamond
coil (see Fig. 133). It is often used in direct-current machines where the end

164 DYNAMO-ELECTRIC MACHINERY

room is limited. It has the further advantage that no coil support is necessary,
as the coils, owing to their shape, fit on to one another, and when banded make
a good strong mechanical construction. It drops further below the iron than
the previous design.

With this type there are as many coils as in the case of the diamond coil,
and the coil itself is the same, with this one exception, instead of forming a
distinct nose where the coil bends over, the change from the upper to the
lower limb is made gradually, this part of the coil having an involute shape.
Further, the upper and lower limbs of the coil are further apart than on the
diamond coil, and no layer of insulation need be put between them. An
armature wound with these ''short" type coils is illustrated in Fig. 133, and
the coil itself is illustrated in Fig. 171.

Fig. 174 shows a mould for a short-type coil. On each side of the figure are
the two halves of the mould separated. Fig. 193 shows how the coil lies in the
mould after it has been wound.

Fig. 194 shows the end of a short-type coil consisting of three sections, and
Fig. 195 gives the construction by which we can determine the dimensions of the
mould upon which a single section of the coil is to be wound.

The involute parts of the short coils are designed so that they all lie together

without binding and without having too much room between them. The involute

curve is described by a point on a string which is being wound on a circle

called the "base" circle. In Fig. 195 the radius of the base circle is r^. The

circumference of the base circle must be equal to tN, where t is the thickness

of the coil and N the number of coils on the armature. Thus the radius of

Nt
the base circle r^ is equal to s— •

This short type of coil must be made in sections, each wound in a mould. The
design worked out below is for the armature coil of a 12 H.P., 220 volt, 350 R.P.M.,
4-pole motor. We are, in the first place, supplied with the data as to diameter and
length of armature, number of slots, etc., given on Design Sheet No. 2.

The first thing to do is to calculate the positions of the two slots in which
the coil lies, then draw the "throw" line a^^ and circle ahc (Fig. 195) along
which the top and bottom halves of the coil touch on leaving the slots. These
things are done in exactly the same way as was described on page 158 with
reference to diamond coils. It is not necessary to draw in more than one of
the sections of a complete coil, because all the sections are the same. Let us
take the one lying on the right-hand bottom comer of slot No. 1 and the
right-hand top comer of slot No. 10.

We have to determine the positions of the points N and M on the throw line
(see Fig. 195), at which normals to the line cut the lower and upper ends of the
involute respectively. The edge a^N in Fig. 195 of the section need not
necessarily lie on the throw line, though in the case we have illustrated we
have made it do so. If it fell below, the wire would, in winding on the mould,
tend to pull down the bevel and make the coil difficult to manufacture. This
remark applies to all 4-pole machines, but the larger the number of poles the
more N can drop below the throw line without causing inconvenience.

THE DESIGN OF ARMATURE GOII£

165

FlO. IM.

Fio. 105. — ^Uethod ot finding the dimensions of a mould for winding a " short-type " armature coll.

166

DYNAMO-ELECTRIC MACHINERY

In fixing the positions of the points M and iV, we have to satisfy the equation

a^N\»Xi e^ = a^M tan 6^ + J" + J",

The Y being the length by which the longer cells has to exceed the shorter, and
the Y ^^ extra for the cutting back of the end and side bevel.

We can make the involute part of the coil longer or shorter as we like. The
further apart M and N lie, the more the coil will drop below the iron, and the

Fio. 196. — Showing construction for finding the centre of the circle nearest to the required involute.

nearer together they are the further the coil will project, until finally, when they
coincide, we get a plain diamond coil.

It will be found in practice that there is nothing to be gained by making
NM greater than shown in Fig. 195, that is, about 0*4 of the throw line.

It is quite unnecessary to shape the mould exactly to the involute curve,
because in any case the coil is flexible and will adapt itself so as to fit in well with
the other coils. It is suflicient to shape the mould to the arc of a circle which lies
most nearly on the involute. A simple way of fixing the position of M and N is
to draw the involute, or the arc that lies near it, on a piec^^ of tracing paper, as
shown in Fig. 196. The circle ahc and the base circle are .the same as in Fig. 195.
Along the circumference of ahc are set ofi^ from M the points e, /, ^, etc., at
distances, giving the pitch of the slots. At centre e, and with radius t equal to
the thickness of the coil, describe the arc of a circle as shown. At centre /, with
radius 2/, describe another small arc. At centre ^, with radius 3/, another, and so
on. The required involute will touch these small arcs. If the involute is to be of

THE DESIGN OF ARMATURE COEUS

167

Tio greater extent than shown in Figs. 195 and 196, the circle whose arc lies
nearest to it may be found very simply as follows. From M draw a tangent
touching the base circle at h From % draw a tangent cutting Mh at B, With k
as centre, draw a small arc from B cutting the base circle. Let the middle point
of this small arc be 0, Then 0 will be found to be the centre of the circle which
almost touches the small arcs drawn from e, /, g^ etc. In practice, therefore,
when dealing with only a small length of involute, we may draw the arc of a circle
at once, finding the centre by the construction given in Fig. 196.

•If the arc drawn on tracing paper is placed over the drawing of the two
sloping limbs of the coil (Fig. 195), and pivotted about the centre of the base
•circle, it is easy to fix ^ and iV so as to satisfy the equation connecting Ma^ and
JVag, and at the same time make the involute long or short, to suit the room
that we have available.

The radius r^ is the radius of the circle formed by the lowest parts of the
coils. This circle must not only be made large enough to clear every part of the
shaft, hub, spider and bearing housing, but should also allow sufficient room for
the circulation of air.

The plan view of the end of the coil is shown in Fig. 194, but need not
actually be drawn.

After the calculations have been made, as given in the calculation sheet below,
we can proceed to fill in the mould designs, sheet No. 2. The dimensions Ay
By C, etc., will be understood from the sketch. Fig. 197.

Dksion Shset No. 2.

MoMld No. 3242.
l8l S.O. No, 62813.
12 H.P. Motor.

Diameter of Armature, 15*.

Length of Armature, 5'.

Number of Slots, 39.

Size of Slots, -407' X 1-5'.

Coils per Slot, 39.

Turns per Coil, 3 sections each of 5 wires.

Size of Wire, "092* d.cc.

Coils in Slots, 1 and 10.

Length of Cells, 6' and 6j'.

Wire Space, -515".

Specification of Aimature Mould.

nth May, 1913.
For Electrical Specification, No. 521.
220 Volts, 350 R.P.M. 4 Poles.

A ej*.

B 8-74*.

L 6-65".

C 2-68".

M 1-4".

D 3-44*'.

-lVO.

E 2-62'.

P 515".

F 1-36".

R\Al.

H^\

8 1:2.

NOTKS.

—Top

leads, 14".

Bottom leads, 13}*.

Depth

of holding

-down wedge, ^.^

Fig. 107.

168

DYNAMO-ELECTRIC MACHINERY

A sketch is made (not necessarily to scale) (see Fig. 198) showing the distances
that the coil projects beyond the end of the iron, and how far it falls below the
working surface of the iron.

Mould No. 3242.

Angle between slots = ^q

Fig. 198.

Calculation Sheet of Short-Type Coil.

10-1

X 360=83°.

Chord on diameter of armature = 16 sin ^= 15 x •062=9-94".

Thickness of coil=3x 103+ •06=0-369".

39 X 0*369

n =

0103 = 8-74,

=2-29,

Therefore

_ ^ ,. 2irx6-64 ,^

Pi at radius ra= — =^— =1'07.

^1=20-2° and tan ^1=0*368,

a,J/x tan^i + •375=2^68 x -368 + ^375 = 1 -36",

2rx4-96
39

Pa *t radius r^=

=0-8.

Therefore

/% t>D9 J/11

8in^2 = -:g-=-461,

6.2 = 27 -5" and tan ^j = -52,
a,.Vxtan<?2=2^62x •52=r36.

FORMERS FOR CONCENTRIC COILS.

Coils which form part of a winding such as illustrated in Figs. 112, 113 and
114 are sometimes spoken of as *' concentric " coils. When open slots are used,
a coil can be inserted after it is wound and insulated, but when semi-closed slots
are used, the coil is only formed at one end, the other end being left open so
that the straight limbs can be pushed through the slot and connected up in position.

Fig. 199 shows a number of concentric coils intended to be placed in open slots.
Both ends of the coils are " bent up." When a two-tier winding is made with coils
that are pushed through semi-closed slots, it is usual to make the ends that are
connected up in position to form the part of the winding that projects straight
out as in Fig. 113a and in Fig. 114.

THE DESIGN OF ARMATURE COIIfi

M).— Two vlewi ol umature coUs o[ tlie " canceatilc " type, made by the OsiUkon

170

DYNAMO-ELECTRIC MACHINERY

In designing a mould foi coils of this type, the fiist step is to lay out the arcs
of the ciiolee upon which the coils will lie and then to set ofi the pitch of the coils
as is done in Fig. 200, The clearances between the coils must be obtained from
the insulation specification, due regard being given to the allowance of space for

5=X

-^ i^

FIS. 200. — Laroot ol thiee conwntrlo ooilB.

air to circulate. From this drawing and from our knowledge of the length of
iron and the length of projection of the cells (see page 172), we proceed to nuke
out the coil winding instractions given in Design Sheet No, 3. These instructions
relate to the coils of a 150 h.f. 3-phase induction motor, wound for 2000 volta,

having 22 poles. There are 9 slots per pole, and 9 wires per slot. The completed
winding is shown in Kg. 114. The bottom portion of the winding instructions gives
a diagrammatic view of the coils. The letters A and B represent lengths which
are specially specified, so that all joints in the wires on the straight end are

In order to settle upon the lengths of wire that will be required for each coil, it
is convenient to have a table such as Table VIII., giving the dimensions of the

THE DESIGN OF ARMATURE COII£

171

TariouB parte of the overhang, lettered A, B, 0 and D, on Kg. 201, The dimension
A is the sparking distance over the surface of the inaulation. It really should
depend upon the voltage to earth, which in three-phase machines is often less than
the voltage between phases. It is, however, good practice to take the volt^je to

Dksign Sheet No. 3

MADHINI DIPT.

COIL WINDING INSTRUCTIONS.

^ } e;i.. —

III tHm\ftmSA^.-.

p«ttun of LHd* . G.cntn3L—

Afeb- /"Si^wa e/-ksii-/4 Cv:

e-Ht^a^ tuHM.

^ji _1;i^"":;;«^^.

.,;„:. ^ .

'<!

/•s

PQ*

+ ■

sS|s*

/i7il

le.

fsn*

r«

/St^

/6.

rw*

8S>b«

/7S

2HA

'_.a

saa

\ '

>«(tn..r»M. 1

affl

^n- 1

Ihep ElH. apM. No. /T*-^?/

earth as if it were the voltage between phases, as this allows for the accidental
earthing of one terminal. In many machines, the insulating tube is put on the
straight part of the coil as in Fig. 202, and the end taped over afterwards. In
cases where this taping can be impregnated so as to make the insulation to earth
over the whole coil strong enough to withstand the full testing pressure, the dimen-
sions A, B, C and D can be considerably reduced. But experience shows that

172

DYNAMO-ELECTRIC MACHINERY

the preservation of these distances is of great service in guarding against accidental
weakness in the insulation of the bent parte of the coil. In calculating the length
of wire regard must be had to the dimensions x and y shown in Fig. 201.

Table VIII. Dimensions of the Ovebhang of

CoNCBNTBio Coils,

Volts between
phases.

0 and 2>.

600

1-6 oms.

•7

1,100

2-6

1-6

2,200

1-6

3,300

6-6

6,000

7-6

6,600

11,000

16,000

FIELD MOULDS.

These require little explanation. As already pointed out, the coil should be
designed so that when insulated it either fits tightly on the pole with no air
pockets between it and the pole, or so that proper provision is made for the
circulation of air between the coil and the pole.

For wire-wound coils (an example of which is given), the mould consists of
four pieces, i,e. two side cheeks made of wood, and a centre piece on which the
coil is wound, consisting of two pieces of wood fitting together at an angle, and
covered on the winding surface with fibre. The centre piece is recessed into the
cheeks, and the whole bolted to the face plate of the winding lathe.

In the example here given, the pole dimensions are 6" x 5^", and the pole
corners are rounded off to a \" radius. Assuming that the coil is wound direct on
the mould, and then insulated afterwards, and that the insulation between pole
and coil is /^" thick, we then get the size of mould as 6^" x SyV* We must,
however, allow some clearance, so that the coil will go on the pole without injury,
and should therefore allow ^^^ at each side, and Y at each end. The finished
dimensions will then be 6\Y x 5^'. The wire space, or space which the coil
itself (uninsulated) can take up radially, must be got from the drawing of the
machine, and is in this case 4|". In getting this dimension from the drawing,
allowance must be made for the fact that the coil will spring when removed from
the mould (in this case about i"), and the insulation on the top and bottom of the
coil being, say, \", the finished depth of coil radially is 5".

The dimension C, or height of the cheek above centre piece, depends on the
number and size of wires used, and in this case would be made 3 J". The radius D
will depend on the size of wire. Design sheet No. 4 gives all the data for a full
mould.

Where the coil to be wound consists of strap on the fiat, the mould need only
have one cheek (fixed to the face plate of the lathe), and a centre piece in which
to wind.

THE DESIGN OF ARMATURE OOILS

173

For coils wound with strap on edge, a bending machine is necessary to form
the comers of the coil, and the coil should be finally shaped on a mould which
must be of iron to withstand the necessary hammering.

Design Shbbt No. 4.

M<ndd

\9t 8.0.

75 K,W. OefL\
H.P, MoUyr J

eTQ/tnt,

Shunt Field Mould.

Superseding Mould

1914.

ph.

For Elec, Spec, . Ins, Spec,

500 Volts, 4 Poles, 750 Ji.P,M, 25 Cycles,

Length of pole (E.), 6'.
Width of pole (F.), 5j".

Size of 1^^' J -081 dec. innd.

Turns of layer, 48,
Number of layers, 34.
Total turns, 1632.
Field Drawing No. 68214.
Wire Space, 4j".

A 6H". D r.

B 5|".
C3i".
Trial Coils.

NOTBS.—

>»4

CHAPTER VIII.

INSULATION.

In one sense the design of the insulation is the mos.t important part of the design
of a dynamo-electric machine. More money has been lost through the breaking
down of insulation in dynamos than through all the other defects in design put
together. There is always a tendency for the designer to improve his copper
space-factor at the risk of leaving just sufficient space for the insulation. But
dearly-bought experience has shown that the insulation should be made as safe a&
possible, even though we may be compelled to limit the machine in other respects
which we may regard as important.

The mere allowing of plenty of room for the insulation is not in itself
sufficient, so much depends upon using the right materials in the right places, and
in supporting them in a manner which experience has shown to be satisfactory.
Even if one type of insulation costs ten times as much as another, it will be found
to pay in the long run if the cheaper method has in it any risk either from the
mechanical weakness or other defect, for in reckoning the cost we must reckon it
as a percentage on the cost of the whole machine, and in estimating the iisk
of breakdown we must take into account the amount of inconvenience to the
user that a breakdown may cause, and the loss of reputation to the manufacturer.
Though only one machine may break down among ten, the maker of that machine
is widely blamed, while only few people hear of the nine machines that stood
up well.

The main qualities of importance in insulating materials are the following:

1. Mechanical qualities.

(a) Mechanical strength in resisting pressure, tension, bending, bruising,.

shock and vibration.

(b) Ease with which material can be worked by being made into sheets,

bent into various forms, moulded, and turned and machined into
various shapes.

2. Dielectric strength.

3. Specific resistance.

4. Property of being unaffected by moisture.

5. Capability of withstanding high temperatures.

INSULATION 175

6. Heat conductivity.

7. Property of resisting oxidization and change after a long period.

8. Specific inductive capacity in service.

1. Mechanical strength, (a) All the good insulators are mechanically weak in
some respects. Those that withstand great compressive stresses are weak in
tension or lack ductility. Ductility is, of course, an important characteristic of
any material which has to withstand mechanical forces; but it is only in the
metals (all of which are conductors) that we find some ductility combined with
great tensile strength. For withstanding pure compressive stresses, mica is as
perfect a material as one could wish for; but if there are any sharp comers
limiting the area under pressure, some bending stresses will be exerted on the
mica, and there being no ductility, the mica may give way. In the same way, the
vitreous and stony insulators, theoretically, can withstand great compressive
stresses, but it is difficult in practice to be sure that these are not combined with
bending stresses which may bring about breakages. None of the insulators can
be relied upon to withstand high tensile stresses. Cotton and linen fabrics are
probably the most reliable in this respect, if not baked too dry or treated in a
maimer which will make them brittle. Cotton and linen fabrics impregnated with
flexible varnishes at moderate temperatures form the most flexible insulators known,
but their flexibility is destroyed when the material becomes dry (see p. 190).

In the table which is given on pp. 176-177, an attempt is made to state the
mechanical qualities of the various insulators and their adaptability for various
uses. The attempt is necessarily incomplete, because so much depends upon
the quality and state of preservation of the material. For instance, treated
cloth, when new, is one of the most flexible insulators known, and is often used in
positions where flexibility is important. Nevertheless, treated cloth, if kept for a
long time at a temperature of 90' C, will become very brittle indeed.

In order to avoid repetition, and to have a convenient method of referring
to the qualities of the materials, we will use certain letters, as given below, to
denote the suitability of any material :

Mechanical qualities when in position.

A. To withstand pressure.

B. To sustain tension.

C. To resist deformation when warm.
I), To withstand bending.

E, To withstand shock and vibration.

F, To withstand abrasion.

Mechanical qualities during mannfactnre.

G, Can be bent in one direction to form angle pieces.
H, Can be bent in two directions to form corner pieces.
/. Can be moulded when hot.

J. Can be moulded in the raw state.
K, Can be machined from the solid.

176

DYNAMO-ELECTRIC MACHINERY

Table IX. The QualitieB

Mica - - - -

Mioanite

Porcelain

Quartz

Marble

Slate . . . -

Lava - - - -

Olafis - - . .

Asbestos

Asbestos slate

Crystallate -

Vulcabeston

Wood boiled in oil

Press-siMihn, fuller ^
board or pure paper/

Do. , with one coat of \
sterling varnish /

Empire cloth

Treated tape

Cotton covering -

Cotton covering and*!
varnish j

Oiled canvas

Leatheroid -

Fibre, white or red

Ebonite

India-rubber

Gutta-percha

Shellac at 28" C. -

Bakelite

Paraffin at 46' C.

Mbchanical Quautibs.

Finished material to rMdat—

A,
Ar
A^
At
At
A3
A^

At
At

At
A,
At

A,
At
A»

At
A^
A,

Bt
B»
Bt
Bt
Bt
Bt

Bt
Bt

Bt
Bt
Bt
B,
B,

Cj (under pressure)

(7, - - F,

C, - - F,

Ct - - F,

Cj — - r 2

Ct - - F,

Ct
Ct
Ct

Ct
Ct

A
A
A

Et
Et

E,
E,

Ft
Ft
Ft

F,
F,
Ft
Ft
Ft

During maaufiaeture
can

o
a
o
•«>

•g

Ot (if thin)

<?.

Ox
Ox
Ox

Ox
Ox
O,

Ox
Ox

Bx
Hx
Hx

Hx
Hx

Ix
Ix

9
o

Jx
Jx
Jx
Jx

Jx
Jx
Jx

8
•8

Ex
Ex
Ex
Et
Et
Et
Ex
Ex
Ex

Dielectric
strength

v^mean' volts

per mm. at 60 <*

(seep. 178).

15,000 to 40,000

10,000 to 25,000

10,000 to 40,000

6,000

3,000

3,000 to 10.000

5,000 to 10,000

3,000

1,000

1,000 to 8,000

1,000 to 4,000

2,000 to 8,000

5,000 to 10,000

20,000

10,000
5,000
3,000

5,000

1,000

10,000

5,000

20,000

to 30,000

to 20,000
to 10,000 \
to 5,009

to 20,000

to 20,000
to 10,000
to 10,000 /
to 30,000
to 20,000
to 20,000
to 20,000
to 25,000

8,000

INSULATION

177

of Insnlating Uateiials.

Specific resistance
•megohms per cm.^

when dry at
25' C.

(see p. 189). ^

Affected by

moisture

or notw

Hoat
conduc-
tivity
at 20* C.
(see p. 221).

Safe teraperature

Resistance to

oxidization and

change with

time.

Specific

inductive

capacity.

5 to 100x108

Not

•00087

500 or more

Very good

6to8

10 to 6,000 X 108

Not

•00029

130, if under pressure

>»

6to8

1 to 1,000x108

Not if vitreous

•007

May crack

4 to5

1 X 10' to infinity

Not

•006

500 or more

4-5

400

-^ected

•002

May crack

Affected

•0022

»»

—

40(0

Affected

•0002

500 or more

5 X 108 to infinity

On surface

•0002

May crack

i»

5 to 10

16x10*

Affected

•00005

500 or more

»>

—

Affected

•0001

»i

—

16x10*

Partly

—

Good

3x10*

Partly

♦>

—

1,000 to W \

•0004

Becomes brittle

•0005

»»

•0006

»>

1,000 to 108

depending

on the

dryneflfl.

Affected

•00035
•00025

•0005

•0002

90
90

it
»»

>»

•0005

»»

•0005

»»

—

*2 to 100x108

Slightly

•0004

\ Are destroyed
1 in the presence
I of air and
J light

2-5

2 to 10 X 108

Slightly

•0004

2-2 to 2^5

25 to 5x108

Not

•0004

3-3 to 4-9

9xlOP

Affected, unless
vitreous

Not

•0006

Softens at 60
200

Very good

3x10^*

Not

•0002

Softens at 50 ^

Melts at 55 \

I Boils at 370 J

1
1

W.M.

178 DYNAMO-ELECTRIC MACHINERY

As these materials possess the above properties to a greater or smaller extent^
we have attached subscript numbers to the letters, to indicate the suitability of
the material for the purpose under consideration. For instance, C^ means that
the material does not soften or deteriorate at all when exposed to warmth. C^
means that it withstands warmth fairly well, C^ means that it withstands warmth
only moderately well.

(b) Most materials which can be moulded into suitable shape when hot, such
as gutta-percha, shellac and bitumen, and petroleum residue, have the draw-
back that they will not resist distorting forces when subjected to a temperature
above 50" or 60' C. Some of them can be usefully employed for the impregnation
of more solid materials. Ebonite can only be used in places where the tem-
perature is low. Bitumen, rubber, shellac and resinous materials are sometimes
mixed with solids, such as asbestos and mica, to form an insulating material
which is mechanically stronger, but inasmuch as these materials can be moulded
when hot, they will give way slowly if put in warm positions. Some of these
can, however, withstand compressive stresses for any length of time.

Within the last few years a new insulating material named Bakelite has been
introduced. This material, which is supplied in the raw state as a thin, varnish-
like liquid, sets under the action of heat and chemical combination to a hard
amber-like substance of great insulating strength and good mechanical qualities.
It can be used for cementing together layers of paper or asbestos, the resulting
product having very fine mechanical qualities, and resisting very well moisture
and fairly high temperatures. Bakelite is a combination of formaldehyde and
phenol; it resists the action of weak acids and alkalies and a temperature
of 250° C, but is affected by strong alkalies. When being heated in the
course of manufacture, it must be subjected to a pressure of about 180 lbs.
per square inch, otherwise gases evolved in the course of setting will cause it
to be spongy.

Where surfaces of a complex shape are to be covered with an insulator, a
common method is to wind cotton or linen tape over them, and treat this tape
in position with Sterling varnish or impregnate it with bitumen, or with one
of the compounds of petroleum residue and bitumen. This mixture of cotton
fabric and insulating compound forms a material having a certain amount of
ductility and ability of resisting tensile and compressive stresses.

Where a material is supplied in the form of sheets, such as the papers,,
it can be bent up into various useful forms possessing good mechanical
qualities.

The insulating materials which can be cut into suitable shapes from solid
blocks are commonly brittle. A good exception to this is hard fibre, which
possesses many good mechanical qualities, but is treacherous as an insulator.

Porcelain and stoneware moulded as a clay and baked at high temperature
can be used in many cases where a moulded material is required to withstand
mechanical forces.

2. Dielectric strength. In Table IX. will be found the voltages which
various materials of 1 mm. in thickness will withstand. No very definite figure
can be ascribed to any particular material, because different conditions in the

INSULATION 179

application of voltage and the slight differences in material give such wide
di£ferences in the result. For materials of perfectly defined composition, and
of crystalline form, such as white mica, we could obtain definite figures for
dielectric strength if the surrounding temperature and form of terminals were
prescribed, and if the time of application of the voltage and other matters
were kept constant; but in other materials such as cellulose (in its various
forms in cotton cloth and paper, treated and untreated), which may have more
or less traces of moisture in their composition, we can hardly expect to get
constant figures for the breakdown voltage.

It is really necessary to enquire into what happens when the material breaks
down. Where a material such as mica, glass or porcelain breaks down instantly
under the application of a very high voltage, it appears aa if the breakdown
were due to disruption of the molecules under the electric stresses. A hole
is pierced through the material, due apparently to the movement of the material
along the path where the electric current has passed. In some cases it appears
as if an explosion had occurred within the material, and for the instant the
forces of cohesion had been inoperative, or had, at least, been overcome by some
other much greater mechanical force. Where, however, the voltage is not
sufficient to bring about this instantaneous disruption of the material, a break-
down may occur due to the heating of the material by electric conduction
through it, and sometimes this heating effect can be produced so quickly as to
seem almost instantaneous in action. The breakdown of the cellulose insulators
is nearly always due to this heating effect. We can in many cases detect the
heating effect by the discoloration of the material if we take off the pressure and
examine the material just before a breakdown occurs. Sometimes a material will
withstand a high pressure for a few seconds and then break down. Upon the
application of the voltage, an electric current (it may be a very minute electric
current) passes through the material. This current causes a slight rise in
temperature, the rise in temperature increases the conductivity, and as the
current increases, the heating increases in greater ratio. With the increase in
temperature, the conductivity still further increases and the temperature rises
more and more until burning sets in, and a puncture occurs. This is what
most commonly happens where treated paper, treated cloth and other cellulose
materials break down. It will be seen that for punctures of this character,
the voltage which must be applied to effect a breakdown is largely dependent
upon the cooling conditions. Taking the material at the temperature of the
surrounding atmosphere, a certain current will flow upon the application of a
certain voltage. If now the cooling conditions are such that the whole of the heat
generated in the material can be conducted away and dissipated without the
temperature rising more than a few degrees, the final value reached by the current
will not be very great. If at any time heat is being generated in the material at
a greater rate than it is being conducted away, the temperature will rise until
a point is reached at which the heat dissipated is equal to the heat generated.
Such a point can of course only be reached if the rate of increase of loss with
temperature is less than the rate of increase of dissipation of heat with tem-
perature.

180

DYNAMO-ELECTRIC MACHINERY

This matter will be more clearly understood by reference to Fig. 206, in
which temperature is plotted as abscissae, and the losses as ordinates. Let
curve W represent the watts converted into heat in a given piece of insulation
when subjected to a certain voltage, and let curve C represent the rate at
which heat is conducted away at different temperatures. The curve C may
be taken for our present purposes as a straight line. The slope of this line
will be steep if the cooling conditions are good, and small if the cooling con-
ditions are bad. Under normal conditions found in practice, the curve C cuts
the curve W at two points, as shown in Fig. 206. The ordinates of curve W
will increase very slowly at low temperatures and more quickly at high tem-
peratures. Let t^ represent the temperature of the surrounding atmosphere ;

Temperature

FIO. 206.

the losses in the insulation are, at that temperature, w^. Under these con-
ditions the temperature will go on rising to t^, at which point the rate of
generation of heat is equal to the rate of dissipation of heat. If from
any accidental cause the temperature should be raised a little above t^y as soon
as the cause is removed it will tend to fall back to tc^^ sa long as the rate
of dissipation of heat is greater than the rate at which heat is being produced.
If, however, the temperature were raised above the point ^3, where the rate at
which heat is being produced is greater than the rate at which heat is being
dissipated, the temperature will go on rising and rising until the material is
burnt and breaks down. The shape of curve W will depend upon the voltage
applied to the insulation, and on the amount of moisture present.

In Fig. 207 are plotted curves which give for different temperatures the loss
per cubic inch in well dried rope paper treated with copal varnish when subjected
to an alternating E.M.F. at a frequency of 50 cycles per second. The curves are

INSULATION

181

plotted from the results of experiments* made on a pad, consisting of treated rope
paper built up to a thickness of ^ inch, and placed in an oven between two copper
plates 9" in diameter, having well-rounded edges. The voltage was applied
between the two copper plates, while the oven could be maintained at
any required temperature by means of an electric heater and a circulating fan.
The losses were measured by means of an electrostatic wattmeter. It was
found that the losses were approximately proportional to the square of the
voltage so long as the temperature was constant, but a higher voltage, if

. e

0 l0 2O3O40S06070S0$0td0a0

Temperature *C

TiQ, 207. — ^WattB lost in well-dried rope paper treated with sterling vamiah when subjected

to an alternating electric pressure at 50 cycles.

continually applied, would produce a higher temperature and that again a
higher loss.

The curve marked 100,000 volts per inch represents the results obtained
from applying 25,000 volts to the i inch pad. If 100,000 volts had been applied
to a 1 inch pad, the centre of the pad would very soon have become hot, and
the losses very much increased on account of the bad conductivity for heat
offered by a thick pad of paper.

What goes on when an excessively high voltage is applied to paper insulation
can be best described by taking an example. Suppose that the conductors in

*C. E. Skinner, '* Energy Loss in Commercial Insulating Materials," Amer. Inet. Mec,
Engineers, vol. 19, p. 1047 (1902).

182 DYNAMO-ELECTRIC MACHINERY

Fig. 227 are insulated with a |- inch thickness of paper or some dielectric having
the characteristics given in Fig. 207. Assume that the iron surrounding the
insulation is maintained at 40*^ C, and that the cooling conditions of the inside
of the insulation are represented by the line C, That is to say, when the
temperature of the inside of the insulation is 32° C. above the temperature of
the iron, the heat passes to the iron at the rate of 1 watt per square inch, or
as the insulation is only one quarter of an inch thick, heat passes to the iron
at the rate of 4 watts per cubic inch of insulation. If we apply 6250 volts
alternating at 50 cycles per second between the conductors and the iron (giving
us 25,000 volts per inch of thickness), we would find that the paper would only
rise in temperature about l^C, because with 1 degree rise under the cooling
conditions assumed, the rate of loss of heat would equal the rate at which heat
was being generated.

With 50,000 volts per inch the temperature would rise about 4*7 as seen
from the point where the line C crosses the curve 50,000. With 75,000 volts
per inch applied, the temperature would rise to 54 degrees or 14 degrees above
the surrounding iron. With 75,000 volts per inch, the temperature would not
rise above 54° C. so long as the iron remained at 40** C. If, however, the
temperature of the iron rose at 50" C. and the cooling conditions were then
represented by the line C\ the temperature of the insulation subjected to a
pressure of 75,000 volts per inch would go on rising until it reached the burning
point, because the curve 75,000 does not cross the curve C\ Similarly, if the
iron were maintained at 40" C. and the voltage raised to 100,000 volts per inch,
the paper would certainly break down in time, because the curve 100,000 does
not cross the line C, We can imagine a case in which the cooling conditions
are very poor indeed, say with a surrounding temperature of 60° C, and a
conductivity so bad as to allow a dissipation of only one quarter of a watt
for 50° C. rise (cooling conditions represented by the line C"\ then the insula-
tion could be broken down by the application of a comparatively small voltage
per inch. In fact, we say that if the heat generated in insulation could be
entirely prevented from getting away, any voltage however small could break
down any insulation however thick.

This matter is very well illustrated by tests* carried out by Mr. E. H. Bayner
at the National Phj^cal Laboratory. Some of these tests were carried out on
insulating tubes made of manilla paper cemented with shellac in an oven in which
the temperature of the air could be measured. The loss on the insulation when
subjected to high alternating stresses was measured by means of an electrostatic
wattmeter. The tubes had an external diameter of 23*5 mms., and the thickness
of the wall was 1'9 mms. For the purpose of the experiment, the tube was about
50 cms. long, and was sealed with a cork at the lower end. A loosely fitting brass
tube sealed also at the lower end was placed inside, and the annular space between
the two was filled with mercury, a great weight of which was by this means avoided.
A thermometer reading to 0*2°, fixed in a cork, registered the temperature of the
air in the inner brass tube. The circumference of the outside of the tube was

^Journal Inst, Elec. Engineers, vol. 40, page 3. Figs. 208, 210 and 211 are reproduced
from Mr. Ra3nier'8 paper.

INSXJLATION 183

71 cme., and a length of 36'6 cms. was covered with tinfoil. This foimed the outer
conductor, which had an area of 260 aq. ctos.

A series of experiments was carried out to investigate more accurately the
effect of (Ganges of temperature, voltage and frequency on material of this nature.
They were all done without moving the specimen, which was kept in the oven
at a steady temperature, generally at about 24'6° C. The experiments were lettered
^ to IF in chronological order.

The upper curves A to 6 (Fig. 208) show the leanlt of a series of experiments
in which fiOOO volte, 60 cycles, was apphed repeatedly to the same specimen. When
first the voltage was applied, the loss in the specimen
was only about 4 watts. This loss was sufficient to
slowly increase the temperature inside the insulation
and increase the conductivity. At the end of thirty
minutes the watts had increased by more than 60 per
cent., and tite rate of increase of temperature was
correspondingly greater. In another fifty-five minutes
the rate of increase of temperature was so great that ^
the loss curve became almost perpendicular, and if the ^
voltage had not been switched ofi, the tube would have
- broken down in the course of a few minutes. This being
then allowed to cool down somewhat, the volt^;e was
applied again. This time the loss followed the curve B.
The quicker rise in the loss was probably due to the
initio temperature of the insulation and the brass tube
inside it. These curves are characteristic of the behaviour
of cellulose insulation when subjected to a high voltage, "'

and it is probable that before a breakdown the loss always increases in the
manner shown, and brings about the burning of the material. In all curves
^ to 0 the rate of the production of heat was greater than the rate of dissipation
■of heat.

Fig. 208 also shows the behaviour of the material under a pressure of 4000
volts. Here the rate of increase of loss with temperature was not greater than
the rate of increase of dissipation of heat with temperature, and the curve for
4000 volts consequently does not go on rising, but shows a tendency to reach the
steady state, where the rate of production of heat is just equal to the rate of dissipa-
tion of heat. At a lower temperature, liS" C, the Iossgh are lower and the curve
is of the same character. At 3000 volts, the loss is reduced, being proportional to
the square of the voltage where the temperature remains constant.

The total volume of the material under test was about 49 cu. cms., and the loss
per cubic centimetre seems to have varied with change of voltE^fe and change of
temperature in the way indicated in Fig. 209. When the material was placed
in an oven maintained at 24-6° C, the cooling conditions are represented fairly
closely by the straight line CO'. At 5000 volts the rate of generation of heat was
for all temperatures greater than the rate of dissipation of heat. The 4000-voit
curve, however, falls below the cooling condition line, so that the material assumed
a steady state at a little over 0-06 watt per ou. cm. The 4600-volt curve almost

184

DYNAMO-ELECTRIC UACHEHERY

touohee the cooling line, but not quite. It was found that at 50 cycles a piesBUtfr
of 4500 appeared at first as if it were going to give steady conditions ; but after
eighty minutes an inflexion appeared in the curve, and then the watts lost rapidly
increased. This is shown in Fig. 210. When, however, the frequency waa dropped

10 Ci

rfes

if.

30 40 SO eo
Tentpero'ticre

from 50 to 47 cycles, the 4500-volt curve was brought down just enough to touch
the cooling line, and the conditions, became steady at 014 watts per cu. cm. A
further reduction of the frequency to 38 cycles made the conditions perfectly
stable, as shown by curve JV. After the conditions had become very nearly steady,
the frequency was raised to 60, and
after a few minutes to 56.

The eSect of changing the sur-
rounding temperature is shown by
Mr. Bayner, With an oven tem-
f perature of 50° C, so great was the

1 increase of the loss that a voltage of

only 2250 gave unstable conditions,
and would have broken down the
tube in less than eighty minutes.
When the voltage applied is such as
to give a curve which very nearly
''"■ ''"■ touches the cooling line, it is found

that a very small change in conditiona makes a very great difference in the
behaviour of the material. As seen from Fig. 210, a change in the frequency

INSULATION 185

from 50 to 63 is safficient to make the great difierence in steepness seen between
curveB J and K. In Fig. 211 is seen the efiect of a slight increase in the voltage.
At 4500 Tolts we have seen the conditions are so near to reaching stability that it
takes seventy minutea to reach the point of inflezioQ of the curve. An increase of
the voltage to 4600 volts makes the curve rise
more qoicUy, and the point of inflexion is
reached in forty minutes.

The dependence of the breakdown voltage
on the cooling conditions is a matter of great a
importance. A paper only '007" thick that \
will withstand 7000 volts when placed be-
tween two cool copper plates will not with-
stand 25,000 volts when 025 inch thick if
the surrounding temperature is high and the
cooling conditions otherwise bad. In one case
the paper withstauds 1,000,000 volts per inch,

and in the Other case it will not withstand 100,000 volts per inch. There are
of course other reasons why thick pieces of insulation do not withstand as high
a voltage per inch as thin pieces. The potential gradient is seldom uniform
in a thick piece of insulation. Very often there are corners producing a steep
potential gradient where the lines of electric force radiate from some edge of
metal, or if there are no comers there are often differences in the specific
inductive capacity or differences in the insulation resistance on different parts
causing undue stress to be thrown on some particular part. If there is a brush-
discharge from a metal comer, this sometimes heats up the insulation and
causes a breakdown.

One of the reasons why mica, whether as micanite or as commonly used
interleaved with paper or cloth, resists such high voltages per inch of thickness
is that the loss occurring in mica when subjected to an altemating pressure
is much smaller than in the case of cellulose.

It should be observed that the losses in a dielectfio when subjected to an
altemating voltage are much greater than when the voltage is steady. This
has sometimes been attributed to Dielectric Hysteresis. But the analogy with
the hysteresis in iron is not complete.* For a stick insulation will not retain
its static strain for an indefinite period after the surface has been discharged.
If we found that it did and that it required the application of a definite voltage
to get rid of the electnti cation, then we would have a perfect analogy. It
appears rather that the loss in a dielectric subjected to an altemating voltage
is purely ohmic. The material, when the pressure is first applied, allows a
dielectric current to pass, by reason of its specific inductive capacity. The
amount of electricity which will flow into a condenser made of paper or other
impure dielectric depends somewhat on the time that the steady pressure is
applied. More electricity will flow into the condenser in two one-thousandths
of a second than will flow into it in one-thousandth, though very little more

186

DYNAMO-ELECTRIC MACHINERY

will flow into it in two seconds than will flow into it in one second. The
material behaves somewhat as if it had ohmic resistance combined with its
specific inductive capacity. It thus comes about, that when we charge and
then discharge the condenser we have suffered a certain ohmic loss. For low
frequencies the ohmic loss per cycle is constant, so that the loss per second
is proportional to the number of cycles per second. But when the frequency
is of the order of 100 cycles, the loss per cycle is less, owing no doubt to the
fact that the condenser has not time to get its full charge. We therefore find
that at high frequencies the loss is not proportional to the frequency.

Curves are sometimes drawn which are intended to show the ratio between
the voltage which will break down a machine in one second and the voltage
which will break it down in two seconds, and so on. It will be seen from the
foregoing that such cur\'^es, even if plotted for different frequencies, different

300 400 500 SOO

Seconds of appUccution of Tkst Voltoffe

Fio. 218. — Eolation between the time of application of pressure and the safe pressure to apply.

initial temperatures and different test pressures would still be very far from
the truth unless they also took into account the cooling conditions of the insu-
lation and its capacity for heat. To provide for all these matters when dealing
with commercial machines would be impossible. Nevertheless, such a cui-ve,
with all its weakness from a theoretical point of view, is better than no curve
at all. For it is clear that it is not fair to apply the same voltage for ten
minutes that we would apply for ten seconds. We may, by prescribing the
conditions not far removed from those obtaining in actual practice, plot a curve
which at least contains some element of truth. If the testing voltage for a
3750-volt machine be taken at twice the working pressure, applied for sixty
seconds, the frequency not above 50 cycles and the temperature of the test
about 65** C, we may take the lower curve given in Fig. 213 as a safe curve
from the manufacturer's point of view. This curve has been arrived at on the
following assumptions. First, it may be assumed that the insulation should be
designed to withstand continuously twice the normal pressure of the machine.
Secondly, we may say that the insulation should be such that we will not

INSULATION 187

have in its weakest part a greater loss than 8 watts per cubic inch of
insulation, even when four times normal pressure is applied to it. For this
loss per cubic inch would heat up the insulation of the weak spot at the
rate of one degree in four seconds. Taking, then, two watts per cubic inch
(a loss easily dissipated) as the permissible loss in the weakest part of the
insulation at twice the working pressure, the upper curve in Fig. 213 gives
us the pressure tests which could be applied with equal safety for the number
of seconds given by the abscissae. This curve agrees with the results found
in practice so far as such irregular results can be made to agree among
themselves. It would be idle to continue such a curve into the region between
0 and 10 seconds, because, if a machine breaks down in the first few seconds,
it is clearly near the danger limit, and there may be so many reasons for this,
such as condensed moisture, broken insulation or what not, that it is unprofitable
to ask what would or would not have happened if the voltage had been applied
for a shorter time.

The lower curve marked " Guaranteed test pressure " has been plotted simply
hy reducing the ordinates of the upper curve in the ratio of 1*5 to 1. In giving
a guarantee we may take any points we like between the two curves according
to the factor of safety that we may choose. The lower curve is particularly
safe for the long duration tests. This is as it should be, because it is not wise
to risk over-heating any weak points there may be in the insulation by a long
application of excessive pressure.

These curves, of course, refer only to the risk of breakdown by over-heating.
Sometimes the breakdown occurs through the air over the surface of the insu-
lation. A breakdown of this kind is commonly preceded by a brush discharge,
and the time of application of the voltage does affect the result, but in a way
far too complex to be expressed on any curve.

The dielectric strength of the insulation on a machine depends very largely
on its dryness ; the presence of moisture increases the losses and the consequent
heating. This matter is dealt with more fully under the next heading.

The test pressure which may be safely applied to a machine is dependent on
the temperature of the insulation. If the insulation is thoroughly dry and at
the same time cold (say 20° C), it will withstand a 10-second voltage test about
o0% higher than if heated up to 70" C. This we can gather from Fig. 206.
Here again our figures can only be rough, and do not at all take into account
breakdowns arising from sparking over surfaces. When the insulation is warm
and dry, there is less tendency for a flash over to occur.

PRESSURE TESTS.

The pressure test which should be applied to the completed machine to ascertain
whether the insulation is strong enough, is a matter upon which a great deal has
been written. The consensus of opinion appears to be that a fairly high-voltage
test for a short space of time, say one minute, ia more satisfactory than a lower
voltage applied for a long time, say one hour, both from the manufacturer's and
the purchaser's point of view. A high voltage will pick out and break down places

188

DYNAMO-ELECTRIC MACHINERY

where the insulation is cracked much better than a lower voltage, however long it
is applied, while the long application of the voltage to a machine which is slightly
damp may spoil insulation which might otherwise get into perfect condition after
a few weeks of service.

The British Electrical and Allied Manufacturers' Association have provisionally
adopted the following tests applied for one minute between the windings and frame
when the apparatus is at normal working temperature. The test should be made
with a pressure of approximately sine wave-form, preferably at the rated frequency
of the apparatus, but in general any frequency between 25 and 100 is satisfactory.

Bated tenninal pressure of circuit.

Test pressure.

Not above 333 volts ....
Above 333 but not above 1500 volts

1500
2250

f»

»»

2250

»♦

1000 volts.

Three times rated voltage with a

minimum of 1500 volts.
4500 volts.
Twice rated voltage.

According to the German Standard Rules also, the test voltage should be applied
for one minute. The voltage to be applied depends upon the rated voltage of the

f2jt>00
11,000

WOO
woo

OfiOO
1000

tooo

5P00
^000

woo
zooo

1.000
0

<B4

t
*

*
^

''•«

'J^-

^
#
^

■'?.

ii 1

1 .

6 8 10 12 /4 16 18 20 22 24

Testing voltcLye in KUovolts

FiQ. 214. — Carres giving the standard testing voltages for machhies of various rated voltages

in Oermany and in America.

machine as indicated by the dotted line in Fig. 214. For machines designed for
a pressure over 7500, the testing voltage is just twice the rated voltage.

INSULATION 189

The rules of the American Institution of Electrical Engineers are a little different
for machines of low voltage. The full line in Pig. 214 gives the testing voltage
for machines of 10 K.w. and over.

3. and 4. Specific resistance and property of being unaffected by moisture.
The specific resistance of all the materials used in the insulation of dynamo-
•electric machines is, when dry, quite high enough for all practical purposes.
When trouble arises from leakage, it is invariably due to the presence of moisture.
The property of being unaffected by moisture is, therefore, one of the most
valuable that an insulator can have. Mica is remarkable in this respect. All the
papers, even when impregnated with varnish, gum or paraffin wax, will absorb
moisture if left for a long time in a damp atmosphere. When damp there is no one
value for the insulation resistance, as this is a function of the applied voltage.
Evershed * has shown that for cellulose insulations the resistance at a pressure of
50 volts is about 3 times the resistance at a pressure of 500 volts. So long as a
machine is in service and has its temperature kept above that of the surrounding
air, its paper keeps dry ; indeed, the tendency is for it to become too dry and
brittle. The impregnation with varnish and gums and the exterior coat of
varnish are very useful in keeping out the moisture during the short intervals
when the machine is not running. If a machine has been 'out in a cold
natmosphere and is reduced to a low temperature, and is then brought into an
atmosphere in which the dew point is higher than the temperature of the
machine, the moisture will collect in drops all over the surface and will pene-
trate to every place that is accessible to the air. When once moisture has got
into a well- varnished armature, it is a rather difficult matter to get it out.
Heating the armature at first merely has the effect of evaporating the moisture
in one part of a coil and driving it into another part. Sometimes the
moisture which is in the pores of the cotton fabric is driven to the surface,
where it is more effective in reducing the insulation resistance. This is seen
from the way that the insulation resistance'"' of a damp machine falls when
it is first warmed up. The following observations were made on the insulation
resistance, measured at 500 volts, of an armature of a 10 K.W. generator which
had been stored for three years. Full-load current was passed through the
■coils to warm them up, and readings were taken of the insulation resistance
at frequent intervals. The insulation resistance, which started at 0*5 megohm,
in the course of one hour fell to ^ J^ of a megohm. In about 3J hours time it
began to rise, and as the moisture was more completely dried out, the insulation
rose to 60 megohms. Even when the resistance has been very much increased
by passing a current through a machine, it must not be supposed that the
drying-out process is complete. It sometimes happens that some parts of the
insulation form a thin layer of dry cellulose, which has such a high resistance
as to make it appear that the whole insulation is good; and it will be found
that, if the current is taken off and the moisture in other parte of the machine

'^See important paper by S. Evershed, **The Characteristics of Insulation Resistance,"
Jour, Inst, mec. Engrn,^ vol. 52, p. 51. Also ** Electrical Conductivity of Press-spahn and
PiUt," Tedeschi, i4rc/«t?/. Elektrot., 1, No. 11, 497, 1913; " Hvgroscopic Susceptibility of
Fibrous Insulating Materials," W. Digby, In«t, Civ, Eng. Proc.\ 183, p. 285, 1910-11.

190 DYNAMO-ELECTRIC MACHINERY

allowed to redistribute itself, the effect of the heating-up process will again
bring down the insulation resistance. If a machine is fairly dry, its insulation
resistance, when cold, will always be much higher than when hot. This effect ia
most commonly due to the way in which the residual moisture distributes itself
in a hot machine, though, apart from this, the insulating resistances of insulating
materials are lower at high temperatures than at low temperatures.

5. Capability of withstanding high temperatures. The only materials used
in the insulation of armature coils which will stand really high temperatures are
mica and asbestos, and as these materials are usually employed in conjunction
with other materials of a more perishable nature, the temperature* which the
machine will withstand continuously is somewhat below 100* C. If it were
commercially possible to insulate a coil both in the slots and on the ends with
mica, and retain the mica in position with an imperishable insulating cement,
it would, no doubt, be possible to run such an armature at 200* C. or more.
Where asbestos is employed, it is usually in positions which do not require very
high insulation. Asbestos, when not well dried out and saturated with some
varnish, has very poor insulating properties, and the varnishes will all perish
if maintained at temperatures much above 100** C. The effect of the temperature
upon the cellulose insulation is very well shown in a paper of experiments made
by Mr. Rayner.t

The materials were tested

(a) Unheated.

(b) After being heated to 75°C.-100*C.

(d) „ „ 125*C.-150*C.

The list of materials tested covers the whole range of insulating materials
commonly used in electric machinery. In general, it is found that materials
such as press-spahn, manilla paper, oiled linen, and varnished tape, which were
flexible at ordinary temperatures, had their flexibility somewhat reduced by being
subjected to a temperature between 75* C. and 100* C. for six weeks or three
months, and became exceedingly brittle when subjected to a temperature between
100* C. and 125* C. for the same period. The brittleness was tested by bending
the material round pins of various diameters. A piece of press-spahn treated with
shellaced varnish 0*34 mm. thick under treatment

(b) broke on being bent round a cylinder f " in diameter.

(c)

»>

})

(d)

>»

It appears from these results, and from the general experience on electrical
machines, that where the temperature is raised a little over 100* C, all the
cellulose materials lose their tough nature, particularly if previously treated
with varnish. We may, therefore, say that 100* C. is the maximum at which

* See page 256 as to permissible temperatures.

t ** Temperature Experiments at National Physical La>x)ratory," ^oitr. Inst, Elec, Engrs.^
vol. 34, page 617.

mSULATION 191

ordinary insulating materials would withstand continuously, and a safe maximum
temperature may be taken as 90* C. for imvamished papers and 80° C. for var-
nished papers. Plain untreated cotton covering will withstand temperatures below
100" C. continuously. Where the cotton is saturated with enamel, so as to make
the coil into a solid block, it appears that even somewhat higher temperatures
do not disintegrate it, although the cold insulation becomes extremely brittle.

6. Heat conductivity. All the heat which is generated in the conductors,
must pass out by conduction through the insulation. This can only occur by
reason of a higher temperature existing in the copper than in the material, be
it air or iron, which surrounds the insulation. The difference in temperature
between the copper and the outside surroundings for a given amount of heat
passing per square inch will depend upon the thickness of insulation and its
heat-conducting properties. In column 5 of the table on page 177 are given the
conductivity of the various insulating materials, expressed in the number of
calories which will pass across one cm. cube of the material in one second for
one degree difference of temperature between opposite faces of the cube. The
figures given are necessarily only approximate, because so much depends upon the
closeness of fibre. Paper which is very highly compressed has much higher heat
conductivity than a paper of loose texture. The heat conductivity also depends
greatly on the temperature. With cellulose insulating materials the conductivity
at 80* C. is five times as great as at 20* C.

The passage of heat through insulating materials is more fully considered in
the chapter on Heat Paths, page 221.

7. Property of resisting oxidation and changes after a long period of service*
This is one of the most desirable properties for insulation to possess, and it is,
unfortunately, not possessed by any materials except those of a stony nature, such
as mica. Deterioration in papers, cotton and varnish, even when not excessively
over-heated, may be due either to (1) desiccation; (2) oxidation; (3) injury from
nitric acid, formed by brush discharge.

(1) If the cellulose insulators are deprived of moisture, they become exceed-
ingly brittle, as shown in the experiments referred to on page 190.

(2) Many of the varnishes, such as Sterling varnish, linseed oil, etc., dry by
a process of oxidization. This oxidization, which will go on slowly at normal
temperatures, occurs much more quickly at temperatures over 90* C. When any
material is treated with Sterling varnish, it is usually dried in an oven until it
is set, but it should be left in a fairly flexible condition. The process of oxidiza-
tion will, however, continue even at ordinary temperatures, and more rapidly at
the temperature of a running machine, so that in time the flexibility can be no
longer relied on. Where, however, a great number of layers are superimposed,
each layer having been painted with varnish which is dried in position, the outer
layers, to a great extent, keep away the air from the inner layers, and these may
maintain their green condition for a number of years.

Where a coil is repeatedly heated and cooled, a "breathing" action goes on
from all its pores; any air lodging in the interstices of the insulation breathes
out when the coil is warmed, and breathes in when the coil is cooled. This
process supplies fresh oxygen for the oxidization of the material. It is practically

192 DYNAMO-ELECTRIC MACHINERY

impossible to prevent this action altogether. The only safeguard is to use in the
insulation a large percetitage of material which is not affected by it.

Foimation of nitric acid. Wherever an electric discharge occurs through
air, nitric oxide is formed, which will readily combine with any moisture present,
to form nitric acid. This will then attack the copper, forming nitrate of copper ;
and if the action goes on to any marked extent, the insulation between the turns
of the coil will break down.

The conditions which bring about a brush discharge from the conductor in the
coils are as follows:

The brush discharge occurs by reason of the voltage which exists between the
conductors in the slots and the iron of the frame. An air space adjacent to the
conductor and lying between it and the iron, as shown in Fig. 164, may be sub-
jected to so great an electric stress that it breaks down. This excessive stress in
the air may be caused either by

(1) An insufficient thickness of the insulating wall.

(2) The nature of the wall being such as to allow a capacity current, or a

conduction current, to flow through part of it and throw an excessive
stress on the remaining insulation, or

(3) The shape of the conductor may be such that the lines of electric stress

radiate from a comparatively small centre, as would occur in the corner
conductors of Fig. 164, and bring about a high potential gradient next
fo the inner conductor.

The condition under which nitric acid is formed in the interstices of coils
has been very fully treated in a paper * by Messrs. A. P. M. Fleming and
R. Johnson.

They show that:

1. This action is rare in machines having a lower voltage than 3500 to
ground.

2. The action only occurs where air pockets are present, and then only
when the voltage across them is high enough to produce a discharge.

3. The gases produced by the discharge in one part may be carried to other
parts of the coil.

4. The action of the products of the discharge (whether these be ozone or
oxides of nitrogen) on the insulation is commonly one of oxidation, and the effects
produced on different materials are :

Untreated cellulose materials have their fibrous structure destroyed.
The oils and gums used in varnishes are subjected to super-oxidation, and

yield organic acids. Linseed oil is readily affected.
Certain asphaltum compounds are very little affected, and paraiiin wax is

quite unaffected.
Mica is unaffected, but the cements used in building it up are attacked.

5. The deterioration of the insulation may occur when no nitric acid is
detected.

* Journal InM. Eler. Engineers^ January, 1911.

INSULATION 193

6. The disintegration of varnished materials is accelerated by the release of
organic acids.

7. Though chemical action may bring about a short circuit between turns, it
does not commonly cause a breakdown of the slot insulation between windings
And ground.

8. If no breathing occurs, the chemical action will cease.

From a consideration of a number of high- voltage machines in which the
•chemical action has been observed, and other cases in which no action has occurred,
Messrs. Fleming and Johnson come to the conclusion that, where the thickness of
the insulating wall between the coil and frame is as much as one mil for every
35 volts, the danger from chemical action is very small, even though there may be
air spaces in the coil and no special precautions taken. They suggest, therefore,
that this thickness of insulation should be adopted wherever possible. In all cases
it is the voltage to earth which must be considered. On an n,000-volt three-
phase star-connected machine, with the star point earthed, the voltage to earth
•cannot rise, under ordinary conditions, higher than 6350 volts; a thickness of
insulation wall of '18" would, therefore, be sufficient on the above consideration.

Where it is impossible to provide insulation of this thickness, breakdowns
from chemical action may still be avoided by adopting special precautions such
AS the following:

(a) The winding should be impregnated.

(b) The conductors, if possible, should be of rectangular section, so as to leave

little air space between themselves, and should be arranged as shown in
Fig. 160 (see page 138) rather than as shown in Fig. 162, so that there
is the lowest possible voltage between turns.

To further safeguard against short-circuit^ the CQiiductors should have their
insulating coverings reinforced on the slot portion of the coil by strips of mica
or other insulation not affected by chemical action. Fig. 165 shows a good method
of carrying out the arrangement of conductors and slot insulation on an 11,000-
volt three-phase star-connected machine.

When the conductors in a coil are so small and numerous that separators
•cannot be used between them, they should ha insulated to ground so heavily a^
to bring the stresses within the safe limit of 35 volts per mil. When such risky
•coils are used in high-voltage machines, surges or other causes of concentration of
potential between turns near the terminals of the machine are very likely to
produce short circuits. Wherever possible, coils of this nature should be avoided
by winding the machine for a low voltage and using step-up transformers.

EXPERIENCE FROM BREAKDOWNS.

The successful designing of insulation must be based on the experience of
previous failures. Certain materials assembled in a certain way have proved to be
unreliable, while other materials or other methods of assembling have been proved
to be good and safe. It is well therefore to look into a few of the most common
<;auses of breakdown. These we will consider under the following headings.

W.M. N

194 DYNAMO-ELECTRIC MACHINERY

Mechanical injury. Accidents sometimes happen to the insulation during
the process of manufacture. A common accident is the cutting through of the
insulation between two copper conductors by the application of excessive pressure
between two hard surfaces. All parts of insulated conductors which may in the
course of manufacture or during running in service be subjected to great pressure
should be specially protected. Thus, in the making of a coil of insulated wire, it
is good practice to give protection to all parts of the wire that may have to bear
more severe treatment than the rest. The first few turns of the first and last
layers should be taped or protected with stout insulation in addition to the cotton
covering. At all points where the wire crosses over to the next layer it should be
protected. Thick wires, and especially square wires, require protection at such
points with very tough insulation such as press-spahn 'OV thick. It is a good
plan to put a rope or a cotton cord of a quarter of an inch diameter on the inside
corners of wire coils. This cord replaces the first or the last turn of the layer of
thick wires, as shown in Fig. 334, or several turns in the case of small wires,,
and gives mechanical protection to the coils, increasing at the same time the
distance between the copper wire and any metal that may come near the corner
of the coil after it is in place.

There are often parts of armature coils that require special protection either
by extra taping or by the insertion of tough pieces of insulation. However small
the risk of abrasion may be at any such point, it is generally worth while to
insert additional insulation if there is room for it, and the cost of insertion is
small. Wherever wires are not tightly held together, and the conditions are such
that a certain amount of relative motion (however small) can occur between them,
we must look upon the place as one of danger, and eliminate it altogether or take
special precautions in the mechanical protection of the insulation.

The end of the straight parts of an armature coil just at the point where it
leaves the slot is a rather vulnerable part, not only on account of the forces which
are sometimes exerted on ic in getting the coil into the slot, but also on account
of the movement of the part which sometimes occurs when running. For this
reason it is good practice to carry an extra layer of insulation between the
conductors at the point even though the room taken up is considerably increased
thereby. Stockings of cotton braiding which can be slipped over the conductor,
and then driawn out so as to make a tight fit, are very convenient for ginng
protection at such places.

Great care must be taken in the selection of materials for the insulation of
conductors on high-speed machines which are subjected to great centrifugal forces.
If cotton fabrics and varnish are used to make an enclosed insulation, this should
be supplemented with tough fuller board to withstand the pressure, and afford
sufficient insulation if the other covering should by any chance be cut through.
Mica is one of the safest materials to insert in cases of very great pressure evenly
distributed. It is not, however, suitable if any sliding motion can occur between
the surfaces which are pressed together. Fig. 212 shows a good way of insulating
armature conductors which are subjected to great centrifugal forces.

The following are the steady pressures per square inch which different
insulating materials will withstand when placed between two flat copper surfaces..

INSULATION

195

One column gives the safe pressure and the other the pressure at which mechanical
breakdown occurs.

Material

Thlckneso.

Safe premure.
Lbii. per sq. In.

Breakdown proesiire.
LbB. per iiq. in.

Calico -

OOo

3,000

15,500

»> * - -

•010

4,000

16,000

Empire cloth

•007

2,000

12.000

Lioen canvas

•010

4,000

18,000

Fibre -

•125

6,000

25,000

Fuller board

•007

6,000

25,000

»♦

•032

6,000

25,000

Mica (pure white)

•010

10,000

40,000

Micanite

•010

8,000

38,000

Mica under test conditions appears to withstand compressive stresses as well
as copper itself. The breaking up of the mica is due to the flowing of the copper
in the direction at right angles to the direction of the applied force. This flowing
of the copper tears the mica up into small pieces. The danger in loading mica
is that it may be subjected to bending stresses due to inequalities in the surface
upon which it is bedded. For this reason the compressive stress is sometimes
kept below 1000 lbs. per sq. in. Sheets of mica between mild steel blocks
withstand 90,000 lbs. per square inch, without showing signs of distress.

In designing an insulation to withstand mechanical forces, it should be
remembered that all the cellulose insulations (paper, cotton and linen fabrics
and the like) become very brittle after being in service for a long time at a high
temperature. The cotton covering of field coils sometimes becomes nothing more
than a dry insulating powder adhering more or less firmly to the wire.

Resistance to shearing stress. The only case of note in which insulating
material is subjected to shearing stress is the case of wedges in the tops of slots.
In direct current turbo-generators, the conductors are commonly retained in the
slots by fibre, fish paper or wooden wedges, and in some cases the shearing
stresses as well as the bending stresses in the wedges is quite considerable. If
wood (hornbeam or beech) is used, the long way of the grain should stretch across
the slot, so that the tendency is to break the wedge across the grain. The
material should be treated with a varnish like Sterling varnish, and dried at a
temperature not exceeding 90*^0. The ultimate shearing stress of fibre or fish
paper may be taken at 8000 lbs. per square inch, of hornbeam 6000 and of beech
3000 lbs. per sq. in. A factor of safety of 10 should be employed on account of
the uncertainty of the fit of the wedge in the slot.

Over-heating. A not uncommon cause of breakdown is over-heating, which may
either destroy the insulation by burning or make it so brittle that a mechanical
breakdown occurs. Such heating is generally due to a cause over which the
insulation designer has no control. But there may be cases where the trouble is
mainly caused by the way in which the insulation is carried out. Cases have
been known in which the current density in the copper has been quite low, and
in which the main part of the armature coils have been comparatively cool during
the running of the machine, and yet certain parts of the armature coils have

196 DYNAMO-ELECTRIC MACHINERY

become hot enough to char the covering, owing to the fact that the insulation
has been put on in numerous layers which enclosed layers of air, thus forming
a very perfect non-conductor. If half the number of layers of insulation had
been used, and care taken to exclude most of the air by sticking the successive
layers together with varnish, the space thus saved would have allowed the air
to circulate between the coils and a bum-out would have been avoided.

Over-heating and charring of insulations has been sometimes caused by the
so-called drying-out process to which a machine is subjected after it has been
standing idle in a damp situation. Current from an independent source is passed
through the windings for the purpose of warming them up and drying them.
Thermometers are placed in certain parts of the coils to see that they are not too
hot, but as the machinery is stationary, and there is no proper circulation of air,
certain other parts of the coils, whose temperature is not indicated, have been
raised to excessive temperatures.

It is rare that over-heating occurs from excessive stresses upon insulation, but
where an excessive voltage test is applied to a machine for a long time, it is
possible for the insulation to be over-heated locally, as described on p. 179.

Trouble from oiL Sometimes oil from the bearings will get upon the armature
coils and weaken the insulation, so that it breaks down. On railway motors and
on machines where there may be difficulty from oil-throwing, the insulation should
be specially enclosed in an oil-proof sheath.

Trouble from voltage breakdown. The cases in which the insulation of an
armature breaks down on the application of high voltage may be divided into
two broad classes — (1) where the puncture occurs through the insulation directly
from conductor to iron, and (2) where the breakdown occurs across a surface.
The first class of breakdown is nearly always due to some mechanical defect.
For instance, a cell of mica and paper may have been crumpled so that the mica
is crushed. Trouble of this kind can only be avoided by careful supervision of
the methods of wrapping and the provision of a sufficient factor of safety in the
number of wraps used.

A common experience in applying a voltage test to a high-voltage machine
is for a breakdown to occur over a wide surface of insulation between the iron
and the ends of the coils. The action is a very complex one, and in many cases
is difficult to account for, a voltage of 16,000 volts travelling over a surface of
perhaps 2 J". The outer surface of the insulation forms a condenser with the copper
conductor, and this condenser is charged and discharged by sparks which creep,
at first only a short distance, along the surface. The air, having become ionized
by these sparks, becomes a conductor, so that the sparks are able to creep further
and further along the surface, until at last the spark jumps direct from coil to
iron. To avoid this action, it is necessary to provide an amount of creeping
surface, specified on page 172, and at the same time to seal, as far as possible,
the ends of the insulating tubes.

Collection of dirt. One of the commonest causes of breakdowns is the accumu-
lation of dirt. Bar-wound machines of low voltage, with certain parts of the
insulation left bare, may stand up well on their original test, but when running
in service, dirt of a more or less conducting character may cause a short-circuit.

INSULATION 197

For this reason, it is best, wherever possible, to use a completely enclosed insula-
tion, which will be sound, however dirty it may be. In cases where the conductors
must be exposed, as, for instance, on commutators, not only must long creeping
distances be allowed from copper to frame, but this surface must be of such a
character that it cannot easily become dirty. An inside surface upon which dirt
will collect and be held by centrifugal forces will not do. The surface, if possible,
should be designed so that the dirt is thrown off it, and, if possible, all such
surfaces should be accessible, so that they can be cleaned.

Breakdown between turns. Breakdown between turns due to mechanical
injury has been considered in a previous paragraph. Sometimes the breakdown
occurs through the voltage between turns being very much in excess of the normal
voltage. This may occur in those coils of high-voltage machines which are nearest
the terminals. When, for instance, an induction motor is suddenly switched on
to the busbars, an electric wave, originating at the switch, passes along each
conductor with enormous rapidity. The steepness of the wave front will depend
upon the capacity of the conductors connected to the switch, the self-induction
of the cables and the manner of making the contact. It is sufficient to say
that, in some cases, the steepness is so great that, as the wave passes into the coil,
it creates an excessive difference of pressure between the ends of each turn, a
pressure of perhaps thousands of volts. If the insulation between turns has only
been designed to stand a low pressure, it may break down under these conditions.
. It is a good plan, on high-voltage machines which are to withstand sudden
switching on, to design the insulation between each individual turn of the first
few coils near the terminals so that they will stand instantaneously the full
voltage of the machine. For this purpose it may be necessary to cut down the
number of turns in these coils to get sufficient room. The voltage between tufns
on field coils under normal running conditions is usually very small, but in
designing the insulation it must be remembered that, if the field current is
suddenly broken or if an explosive arc occurs in the armature current, a very
much higher voltage may be thrown on each turn. It is therefore well, on field
coils, to be not too sparing with the insulation between successive layers, and the
insulation to ground must be capable of standing 3000 or 4000 volts on field coils
excited at 125 volts and of withstanding proportionally higher voltages in cases
where the exciting voltage is higher.

METHOD OF INSULATING CX)ILS.

This subject is such a very large one that it requires a book to itself, so
nothing more will be attempted here than a statement of the salient points.
Messrs. Fleming and Johnson in their book* have treated of the matter very
fully both theoretically and practically, and given numerous diagrams of shop
methods. For further information the reader is referred to that book and to
other treatises! which specialize upon the subject.

* InsulcUion and Design of Electrical Windings, (Longmans.)

t Turner and Hobart, Insulation of Electrical Machines, (Whittaker. )

198

DYNAMO-ELECTRIC MACHINERY

The insulation of the different parte of a machine may be considered under the
following headings:

(1) The insulation and assembly of the separate conductors.

(2) The insulation of the slot.

(3) The insulation of the end windings.

(4) The insulation and support of the terminals.

(I) The insulation and assembly of the separate conductors. The cotton
covering which has been for so long used to insulate the separate turns of a coil
from one another still seems the best and cheapest material for wires not very
small (say greater than 0*05" in diameter). For very small wires the cotton
covering takes up a great deal of room. Silk covering is rather expensive, but
for small wires it pays for itself by the economy that can be effected in the size
of the machine. Enamelled wire has largely come into use for small sizes, and as
the price of this wire will probably be much lower in the future, one may expect
it to be very commonly used. The thickness of enamel only adds 001* to the
diameter of a 'OV wire, as against -0015" for single silk, -0025" for double silk,
•005" for single cotton, and 009" for double cotton. Enamelled wire is generally
wound with very thin paper (-003") between layers. This paper takes up a good
deal of room, so that the saving in space with enamelled wire is not as great as
it otherwise would be. The voltages (50 cycle alternating) which two round
wires '05 in diameter at a temperature of 20' C. can withstand for one minute
when pressed together with a force of 10 lbs. per linear inch are given below :

Untreated cotton covering -
Paraffined ,, ...

Impregnated (petroleum residue)
Varnish (Sterling) ....

„ (Shellac) . . . .
Untreated silk covering
Paraffined ,, . . .

Varnish (Sterling) ....

„ (Shellac) . - . .

Enamelled wire

Enamelled wire and one piece of *003

paper impregnated with gum -
Aluminium with normal oxide coating

Total thicknem

of insulation

between two wires.

Puncturing voltage.

0-012* thick

1250

•01-r

2000

•012"

2000

•012"

2000

•012"

1500

•0035"

500

•0035"

1500

•0035"

1500

•0035"

900

•001"

300

•005"

1300

less than '001'

/
- -

200

The voltage between adjacent wires in field coils and small armature coils is
usually very small, so that the most important quality of the insulation is that
it shall resist abrasion during the process of manufacture, and shall not deteriorate
on the running machine. The heat conductivity of the assembled wires is also
of great importance, especially in the case of thick coils with a great number of
layers. This matter is dealt with fully on p. 221.

The space occupied by a number of round insulated wires depends very largely
on the bedding of the layers. When the wires are small, and particularly if the
coil is of rectangular shape, one cannot be sure that the turns of one layer will be

INSULATION 199

in the hollows left by the last. Where the wires are of fair size (O'l" diameter),
•and even for smaller wires if the coil is of cylindrical shape, one can, by the
exercise of a reasonable amount of skill, rely on the theoretical amount of bedding
being obtained within a few per cent. In Fig. 166 we give the space factor (that
is the ratio of cross-section of copper to cross-section of winding space) to be
found in coils with ordinary skill with different sizes of wire. With a coil made
to fit a rectangular pole, it is much more difficult to make the wires bed into the
grooves. Some makers use extra insulation between layers at the comers, and
this prevents bedding altogether. Such coils often have a space factor as low
as -5 when wound with d.d.c. wire '05".

It should be noticed that when one layer is wound over another, there must
be a point somewhere in each turn where the wire crosses the wire lying under it.
If this crossing over is done regularly at one side of a coil only, and occupies only
A short length of the wire, the bedding may be good for all the rest of the coil,
but if it is done irregularly, it leads to bad bedding throughout the coil.

Annature coils. The insulation between turns of armature coils must be
•carried out with the very greatest care, as the short circuiting of a single turn
is very disastrous, whereas the short circuiting of a turn on a field coil is of
•comparatively small importance. Double cotton-covered wire is commonly used
for armature coils where the size of wire is not greater than No. 12 s.w.G.,
And where the voltage between turns is not greater than 25. Double cotton
covering is preferred for armature coils because it is not so liable as single
cotton to open out and show bare copper at places where the wire is bent
■around a corner of small radius. At parts of a coil winding where abrasion
may possibly occur during the* winding or during the operation of the machine,
it is advisable to supplement the cotton covering by a layer of tape or tough
paper. In the case of square or rectangular wire, a very tough paper or leatheroid
■should be used between turns at cross-overs and at all points where the corner
of the rectangular wire may bear with undue stress upon the cotton covering.
Cotton covering is much improved, both mechanically and electrically, by being
treated with a tough varnish. If the wire is treated before winding, it should
not be dried so hard as to cause the covering to crack when wound around
■comers. Armature coils and other windings of no great bulk are commonly
dipped, after being thoroughly dried, in a copal varnish, and the outer layer of
varnish is oxidized in an oven and forms a coating which is fairly efficient in
keeping out moisture!. When a coil is very bulky, it is impossible to
oxidize the varnish on the inner part, so it is better practice to impregnate
coils of considerable section in a vacuum oven, and force the impregnating
compound into the interstices of the coil by the application of external pressure.
For larger wires, the insulation should be supplemented with a wrapping of
tape. Large round wires are now seldom employed in armature coils, wire of
rectangular section or strap affording a much better space factor and giving
better mechanical support. Double cotton covering is used on very small straps,
but an insulation of mica and paper is to be preferred on the straight part of the
coil. The common method of insulating the straps of a direct-current armature
coil to stand up to 500 volts is to interleave one end of a sheet of mica paper

200 DYNAMO-ELECTBIC MACHINERY

between the straps, and then to wrap the remainder of the sheet 2J or 3| times
round the conductor as a whole. This, of course, c^ti only be done on the
straight part of the coil. In the diamoud-ehaped or involute ends of the
coils, each strap must be separately taped or alternate straps may be taped
with an overlapping layer, and the whole coil treated before the mica-paper
insulation is applied. The tape on the ends must be continued for |" into the
straight cell, and tor this reason the straight cell is made to project 5" beyond
the slot. This portion of an armature coil, just at the ends of the slots, has given
much trouble in the past, and requires very careful treatment. If two coils lie
one above the other, it is desirable that the cells of both coils should not end at
the same point. For this reason, one of the cells is made to project a full inch,
as shown in Fig. 198.

FlO. ZIZ.— Showa the metliod ol taping the Individual turns of a coacsDtnc winding.

In A.C. generators, the voltage between adjacent conductors is Bometimes of the
order of one or two hundred volts, and at instants of switching on may amount
to several thousands of volts. In these cases it is well to have a really good
mechanical separation between the conductors, such as a stiip of mica or mica
paper, and each conductor should be separately taped with overlapping layers.
Fig. 212 shows the method of taping the individual turns of a concentric winding.
After assembling, the insulation at each successive turn should be tested at
3000 volte, where the voltage per turn is not more than 75 or 100 volts, and
5000 volts between tunis if the voltage is over 100 volts per turn. It was stated
on p. 197 that it is well to insulate adjacent layers of the first and last coils in a
high-voltage machine, so as to withstand for an instant the full pressure of the
generator between turns.

(2) The insnUtion of the coil from the slot. The most relial)le material for the
insulation of the coil on the straight part is mica, because it resists damp and does
not undergo any change with time. A cell or tube of mica built up of shellac is.

INSULATION 20r

however, rather too brittle, and it is therefore better to use some paper in the-
composition of the cell. Some makers use mica and empire cloth. This gives a
cell rather more flexible than the pure mica,* and if the portion of the coils which-
projects through the slot is accidentally bent through a small angle, the paper or-
empire cloth affords a cushion to the mica, and prevents it from cracking. The
manufacture of these wrappings has undergone a long course of evolution, and
a great deal of experience and skill is necessary to produce a really well-wrapped
coil free from avoidable air spaces and yet of sufficient flexibility. The thickness-
of mica and paper wrapping to be employed may be taken from the data given by
Messrs. Fleming & Johnson in their paper above referred to. In no case should,
the thickness be less than '04". Having provided sufficient mechanical strength,
the provision of xwujf' ^^ insulation for every 35 volts above earth has been foundi
to be sufficient in actual running machines. Where the space available on a
machine does not permit of the full thickness according to this rule, then special
precautions must be taken, as stated on p. 193, to prevent injury due to-
brush discharge.

There are two general methods of applying the external insulation to an armature
coil. (1) The insulation may be wrapped around the conductors and completely
finished before being put on the machine. (2) The slot in which the coil is to lie
may be lined with an insulating tube or trough before the conductors are inserted.

The first method is to be preferred for high-voltage machines, because it enables
the coil to be properly sealed and made into a complete whole, able to withstand
a high-flash test and prevent the formation of nitric acid (see page 192). The
second method must be used in those cases where individual conductors must
be inserted in the slot one by one. The mush- wound coils illustrated on page 122,
the field coils illustrated in Fig. 133, and all hand-wound coils are instances where
the slot must be lined with insulation before the insertion of the conductors. In
all such cases it is desirable, where the carcase is not too big, to place the whole
machine when completely wound in a vacuum oven, and, after exhausting all
moisture, to impregnate the windings under pressure. This has the effect of sealing
the overlapping layers of insulation at the openings of the slots.

EXAMPLES OF CALCULATIONS OF ROOM TAKEN BY INSULATION

OF ARMATURE COILS.

Mush-woand coils for yoltages up to 600. The slot may be insulated with one
piece of press-spahn 0 05 cm. thick, and one piece of varnished cloth, such as " Empire
Cloth," of a thickness of 0*02 cm. After the wires are inserted, the cloth is folded
down so as to overlap. A suitably shaped tool is then pushed between the press-spahn
and the overhanging tops of the teeth to press the edges of the slot lining well down
and enable a strip of leatheroid or press-spahn about 0*08 cm. thick, and almost
as wide as the slot, to be inserted as a lid to the slot. The parts of the coils lying
outside the slot may be taped with " Empire " tape, and then with ordinary cotton

*** Experiments with Paper-free Mica Tubes," K. Fischer, Eltktrotech, Zeitachr. 31,
p. 239, 1910 ; *• Artificial Mica," EUctrochem. Ind,j N. Y., 6, p. 257, 1908 ; " Use of Mica as an
Insulator," F. Wiggins, Elect. Rev,, 71, 564, 1912.

1202

DYNAMO-ELECTRIC MACHINERY

tape. After the coils are connected, the whole is impregnated in a vacuum tank.
Mush windings of this kind may be wound either with one coil per slot or with
two coils per slot. In the latter case, the varnish cloth covering must be lapped
•over the lower coil and a spacer 0-05 cm. thick inserted before the cloth insulation
of the upper coil is inserted. Special taping of the parts of the coils lying outside
the slots is necessary to prevent coils of opposite polarity from coming in contact
with one another.

To arrive at the amount of space required for the insulation of continuous-
<^urrent armatures for voltages up to 600, we can make the calculation as follows :
The paper will be about 0*013 cm. in thickness, and the mica may make up the
total thickness to 0025 cm. The curved part of the coils lying outside the slot
may have alternate straps taped with 0^015 cm. tape, the whole coil being dipped
and dried before the straight parts are insulated. Mica tape on the ends is some-
times used where there is some fear of the coils being subjected to a high temperature
or exposed to moisture. The external insulation of such a coil may consist of
2^ turns of mica and paper or mica and cloth, each turn being about 0-025 cm.
thick. The whole coil is then taped over with cotton tape, which is half lapped
on the projecting ends and wound without overlapping on the slot portions. A
slot lining of 0 02 cm. paper is generally used with coils of this kind, making the
total thickness of external insulation about 0-1 cm. from copper to iron. The
width of slot in cms. required is therefore

m«K+0 025w+0-2+/r,

where m is the number of straps side by side, tg the thickness of the straps, and
fg is the allowance made for roughness inside the slot, usually 0 05 cm. to 0-07 cm.
Between the upper and lower coils lying in the same slot, it is well to place a piece
o{ 0*050 press-spahn, so that the taped portions of the coil which are of opposite
polarity may not be too near together. It is also well to place a liner at the base
of the slot, which may be of 0025 press-spahn.

Table X. Allowance of Room in Slot fob the Extebnal Wbappino
OF Abhatube Coils of A.C. Gekebatobs and Motobs.

Length of Iron np to

Length of iron up to

Length of iron above

Voltage of machine.

80 cms.

100 cms.

In width. In depth.

Tn width.

In depth.
•47 om.

In width.

In depth.

2,000

•26 oxn. -36 om.

•35 oxn.

•45 cm.

•58 om.

4,000

•32 -42

•42

•54

•62

•67

6,000

•4 -5

1 47 ' 59

•68

73 '

8,000

•45 -55

•53 -65 I

•62

•77

10,000

•5 6

•58

•7

•68

•83

12,000

•6 68

•65

•76

•72

•87

14,000

•66 -76

•75 1 -87

•85

1-0

The stator coils of A.O. generators and motors. Where the conductors are
small, double-cotton covering is used. For this allow a thickness of 015 cm.,
making the space occupied by each conductor -03 cm. wider and deeper. Where
two conductors are in parallel, as in Fig. 161, the space occupied by the double

INSULATION 203

•cotton covering in width will be 06 cm. Where the arrangement is as in Fig. 162,
•an allowance of -08 cm. should be made for each strip of press-spahn. Where the
•conductors are taped individuaUy, the thickness of tape is generally -04 cm. radiaUy,
^ving a total extra width of *08 cm. In high- voltage machines a strip of mica
will generally be placed under the tape about -06 cm. thick. The room to allow
for external insulation can be taken from Table X.

(3) Insulation of the end windings. The distance which a straight cell should
project beyond the slot for different voltages is given in Table VIII. page 172.

For all voltages over 3500 it is very desirable to have the insulation com-
pletely sealed at the ends of the cells in order to avoid the creeping action
described on p. 196. Where open cells are used, the coil can be impregnated
^as a whole, and a well-sealed insulation obtained.

Where end connectors are employed, these should be individually taped with
overlapping layers of tape or Empire cloth, the whole being treated with varnish,
re-taped, and treated several times in succession until the connector will withstand,
when in position, the full voltage of the machine. In assembling, the connectors
are separated by press-spahn at all parts where they come under clamps, and
treated wooden blocks are used to preserve a good sparking distance to earth.

(4) The insulation and support of the tenninals. The conductors which lead
from the winding to the terminals of the machine are usually insulated with
successive layers of treated tape of such thickness as to withstand the full test
pressure of the machine.

In bringing out the terminals of a high-voltage machine, the greatest care
must be exercised. Only materials which retain their mechanical qualities
should be used. Rubber-covered cable is not recommended, as it may soften
with the heat or become satmrated with oil. The terminal conductors must
be held firmly in position by strong clamps, so that the insulation of these con-
ductors must be of a kind that can resist mechanical pressure for any length
of time. Some makers attach the conductors to porcelain insulators. This
if carried out on a very substantial manner is good, but even the strongest porcelain
insulators are sometimes broken in shipment. Another plan is to wrap the terminal
conductors with a great number of layers of varnished cloth, each layer being
treated with Sterling varnish before the next is applied. In this way a very tough
and strong insulation can be built up, which entirely closes the conductor in and
which can be clamped between wooden cleats.

When large generators and motors are installed, it is seldom worth while to
provide terminaLs to the conductors which can readily be connected and discon-
nected, because one connection is all that is generally necessary in the lifetime of
the machine. It is suj£cient in general to provide thimbles by which a permanently
sweated connection can be made. The thimble should by preference secure the
cable without relying on the solder for mechanical support.

CHAPTER IX.

VENTILATION.

The high electrical outputs which, in modern machines, are obtained from com-
paratively small amounts of material are due mainly to the improvements that
have been made in methods of ventilation. The subject is therefore one of th&
greatest importance from a commercial point of view. The tendency now is to
design a machine as we would design a blower,* providing definite paths for the air
as it comes in, an efficient means of blowing, and a definite path for the air to the
point where it is expelled.

Before proceeding to consider the various systems of forced ventilation, we will
take up a few important matters relating to self-yentilating machines, that is to
say, machines through which the draught is produced by no other means than the
rotation of the working parts.

The first point is that the general shape of the frame and of the rotating part
should be such that the warm air is thrown far away from the machine. The
tendency for hot air to rise is not always sufficient to take it away from the
neighbourhood of the machine, and it sometimes happens that the same air is-
drawn into the machine again and again, thus causing a very much higher tem-
perature than would be obtained if the main supply of air were at the temperature
of the room. A common cause of this trouble is the shape of the end bells or
overhanging frame, which gives to the expelled air a horizontal direction and
carries it to the vicinity of the intake. In self-ventilating machines, we ought to-
see that the centrifugal blowing action of the rotating parts is allowed to give the
air a yelocity which takes it well away from the intake. Sometimes a continuous-
current generator which will run fairly cool when running by itself will have an
excessive temperature rise when direct coupled to a motor on account of the
interference with its scheme of ventilation. A little forethought and suitable
shaping of the parts will obviate difficulties of this kind. With motor-generators-

* The following references will be useful to the reader : "Turbo-generators and High-speed
Motors," Niethammer, Ztitnchr. Veriines Dtut^ch. Imj., ,53, pp. 1009, 1313, 1406, 1909;
"Coohng Ducts, Use of in Electrical Machines," T. Hoock, Ehk. \i. MaschinenhaUj 28, p. 908,
1910; "Ventilation of High-speed Dvnanios," K. Czeijii, Efektrotfch. ZeiUchr.ydZ^ pp. 313 and
343, 1912; "Ventilation of Turbo-alternators," K. Knowlton, Electrician, 70, p. 259. 1912;
"Ventilation Arrangement for Large Generators," Weltzl, ElelU. u. Mtuchinen^u^ 31, p. 10,
1913 ; " Air- filtration. Cooling and Ventilation of Electrical Machinery'," Christie, JSUclriciafiy
71, p. 452, 1913.

VENTILATION

206

•a

206 DYNAMO-ELECTRIC MACHINERY

it is a good plan to arrange the fanning action so that the air is blown out radialljr
between the two machines and drawn in at the ends of the set. Sometimes the
flywheel adjacent to a generator direct coupled to an engine will prevent a proper
supply of air to the electrical machine, while a very little variatioii in the dis-
position of the parts might make the flywheel improve the ventilation. Some-
times a machine, when running by itself, will be fairly cool, but when adjacent
machines are running it gets hot from the air thrown off by them. It is therefore
necessary, when checking the actual temperature rise with the rise expected from
the calculation of the machine, to see that there are no abnormal external circum-
stances which make the temperatures either higher or lower than they would be
on a fair test. For instance, if a dynamo intended to be direct coupled to an
engine is on test driven by a belt, the windage from the belt will sometimes keep
down the temperature rise by several degrees.

Amount of air required. Sufficient air must be provided to carry away the
heat generated. A supply of 100 cub. ft. of air per minute for each kilowatt
loss will in general be sufficient. If the conductivity for heat of all parts is.
sufficiently good and the air is so evenly distributed that none of it receives a
temperature rise greater than 32** C, it may be that 60 cub. ft. of air per minute
would be sufficient to keep the machine below 45° C. rise.

Having provided a supply of cool air to the machine as a whole, the next step
in the ventilation problem is to see that the openings in the spider are sufficient
to carry the air to the ventilating ducts. One of the reasons why machines of
short axial length come out in practice to be more economical than would be
expected from a calculation of the amount of material they contain is that the
supply of air from the two ends, not only to the end windings, but to the
ventilating ducts, is much better than on machines of greater axial length and
smaller diameter.

The ventilating ducts themselves must not only be of sufficient cross-section
to allow enough air to pass ; they must also present sufficient cooling surface for
the heat to pass from the iron or copper to the air. In the next chapter we give
specific figures for the amount of air required and for the rate of passage of heat
from the various surfaces. In this chapter we are only concerned with the general
schemes of ventilation.

Schemes of ventilation. Fig. 215 illustrates a scheme of ventilation commonly
met with in turbo-generators. Here centrifugal blowers are placed at each end
of the rotor, and supply air to the completely enclosed ends of the machine. The
air, after blowing over the armature coils, finds its way partly along the air-gap
and partly through axial ducts in the rotor, from which it is thrown out by the
radial ducts. The air then passes through the radial ducts in the stator iron to
the annular space in the frame, and is finally expelled at the top of the machine.

When a machine is of great axial length, it is sometimes not possible to get
enough air along the axial holes in the rotor, and for this reason other methods
must be adopted. One method, illustrated in Fig. 216, still employs radial ducts,
but the air is caused to flow inwards radially in some sectors of the machine, and
outwards radially in others. The figure is self-explanatory. It will be seen that
not only is there practically no limit to the amount of air which can be supplied

VENTILATION

20T

«
«

o
C

^08 DYNAMO-ELECTRIC MACHINERY

to the middle of the machine, but the air coming direct from the fan is cooler than
air which has passed through the rotor.

Radial ducts and axial ducts. The cooling surface necessary for the passage
of the heat from the iron to the air may be provided either by means of radial
ventilating ducts, as illustrated in Figs. 215 and 367, or by means of axial ducts,
•as shown in Figs. 218 and 220.

Eadial ducts are made by placing "ventilating plates" at frequent intervals
l>etween the ordinary punchings, and are convenient in design, in so far as the
number of them can be easily altered to suit the circumstances of each case
without any interference with standard punchings. In the rotating part they act
SLS blowers, drawing their own air in machines that have no separate blower, and
supplementing the special blower when one is provided. Fig. 429 shows the
scheme of ventilation of a 75 K.w. D.c. generator fitted with a blower at the end
opposite the commutator. In this machine the rear-end casting is formed so that
it converts the rotational motion of the air into an outward blast whichever way
round the machine is run, and thus the fan acts as a fairly efficient blower, causing
the air to enter at the commutator end. Part of the air is drawn through the
-channels in the armature and part is drawn between the field coils. The space
between the fan and the rear end of the armature is contracted and throttles the
flow somewhat at this point, so that while a sufficient amount of air is drawn
through the armature to feed the ventilating ducts, the blowing action of these
<iucts is not overpowered by the sucking of the fan. It will be seen from the
•calculation of this machine given on page 489 that the cooling coefficients of the
field coils and armature coils are greatly increased by the use of the fan. A
:somewhat similar system of ventilation is illustrated in Fig. 217, but here the
spider is completely closed at the slip-ring end. The air drawn in by the fan
and ventilating ducts is driven through the stator to the opposite end of the
machine.

In Fig. 218 is given the scheme of ventilation of a railway motor, in which
the fan is placed at the commutator end, and, instead of radial ducts, we have
^xial holes running through the armature.

Axial yentilating ducts do away with the necessity for ventilating plates, and
thus enable the straight part of the armature coils to be made somewhat shorter
in those cases where sufficient iron can be obtained behind the slot without
lengthening the frame to make up for the -space occupied by the holes in the
punchings. The conduction of the heat takes place very much more readily along
the direction of lamination than across the laminations (see page 251), so that the
heat travek through the iron to the surface of axial ducts more readily than it
-does to the surface of radial ducts. In practice, however, radial ducts are placed
at fairly frequent inter^'als, so that the drop in temperature between the centre
And edge of a packet is not of very great importance (see page 390 and Fig. 367).
Axial ventilating holes should not be made too small, as the rough surface
presented by the edges of the punchings renders them rather liable to be stopped
up by dirt. A diameter of 30 millimetres or more is usual. The holes being
fairly large in diameter, the ratio of the area of duct surface for a given length to
area of cross-section is much smaller than for radial ducts. For this reason one

VENTILATION

209

•a

a
«>

-8

•S

■*>

§
I

tcod

2
•"I

o
to

■c

A
•

W.l#

210

DYNAMO-ELECTRIC MACHINERY

would expect air to travel for a greater distance along an axial duct before it
attained ita full temperature rise. The presence of the holes In the punchings
interferes with the magnetic circuit, so that either a much greater depth of iron
must be used or the total length of iron in the machine must be increased. When
the depth of the punchinga is increased there is no limit to the amount of air
that can be supplied to the centre of a machine by means of axial ducts. This ia
of great importance in the design of very large turbo-generators.

Fig. 219 shows a scheme for ventilating a turbo-generator in which air
is blown from both ends of the machine through axial ducts to radial passages
near the centre of the armature iron, whence it escapes into the annular space
around the frame. This method is suitable for machines that have great axial
length. Where the axial length is not too great it is sufficient to blow the air
from one end of the machine only.

The method of axial ventilation, in which the air is passed from one end of the
machine to the other, is well illustrated in Fig. 220. The rotor is of the type
shown in Fig. 361, and has no radial ducts. Below the space in each slot provided
for the winding there is a channel for carrying air. The air thus passes close
to the place where the heat is produced, and by cooling the root of the teeth
enables the heat to pass readily through the insulation of the coils in the slots.
The blower at one end of the rotor (the left-hand side in Fig. 361 and the
right-h«id side in Fig. 220) consists of two parts. The inner part throws out

VENTILATION 211

the air from the closed end-bell and cauees it to be drawn in from the opposite
end of the rotor along the ducts immediately below the conductors. The other
part of the blower supplies air for cooling the armature conductors at that
end of the machine. On the other end of the rotor (the right-hand side in
Fig. 36! and the left-hand side in Fig. 220) a specially wide blower is provided,
which supplies the air to cool the armature coils at that end and also to cool the
armature iron. The air, after being forced into the enclosed end-bell of the stator
(where it is received by suitably shaped surfaces which convert its tangential
velocity into pressure), passes along numerous axial holes in the stator iron to the
other end. It then passes through holes in the stator frame into tbe annular
apace behind the iron, from whence it is conducted to a flume at the base of the
machine. The passing of the air through the machine from one end to the other

Fw. 219.— Schemt

UatlDii ol tuibo-fleuer&tor by meuis ot bolag la the sUmpinis parallel
m : klr pualDg from both ends ol mmchlne (Meeara. Siemens).

will of course cause one end to be hotter than the other; but there is no serious
disadvantage in this, provided both ends are cool enough. The fact that only
warm air is provided for cooling the inside surfaces of the rotor conductors at one
end must, however, considerably reduce the rating of the machine. Where the
turbo-generator is very long, it is better to pass the air through the iron from both
ends to the middle, or to adopt the method illustrated iu Fig. 216.

It will be seen on page 242 that the value of hv (the watts per sq. cm. per degree
C. difference of temperature between surface and air) is dependent upon the v, and
as it is the velocity of the air iu intimate contact with the surface that is of chief
importance, we may gather that for a given quantity of air passed through the machine
narrow ducts will be more effective than wide holes. The ducts, however, must
not be too narrow or they will be liable to be stopped up by the accumulation of
dirt. If a duct is too wide the air passes through it without taking as much beat
from the iron as it would if it were passed through a narrower duct. Experiments

212

DYNAMO-ELECTRIC MACHINERY

VENTILATION

213

have been made from time to time to determine what is the best width of duct,
and it seems to be generally agreed that for large machines having great depth
of iron the air ducts should be about 1 cm. wide. If, however, the pressure
available for forcing the air through is high enough, and especially if the air is
filtered so that there is not so much danger from dirt, it seems to be better to
choose a rather narrow duct. It will be seen that in the 15,000 K.v.A. machine
illustrated in Fig. 367, we have ducts only ^" wide. As the machine is ventilated
by means of filtered air from an independent blower at a considerable pressure, rather
narrow ducts can be used. We are thus able to use a very large number of ducts
without taking up too much room, and these present an enormous cooling surface.

Yet another method which is very eiFective for long turbo-generators is
illustrated in Figs. 221 and 221a. There the ducts are made much like radial ducts
with ventilating plates ; but the spacers in the plates are not radial. They consist
of concentric ribs which allow the air to enter at the top of the machine and go
out at the bottom. The supply of air may either come from fans in the rotor
shaft, aa shown in Figs. 215 and 220, or from an independent blower. The air
that cools the rotor is drawn through channels in the shaft by means of the
centrifugal action of the ventilating ducts, and is expell^ into the air-gap, from
whence it passes through a certain section of ventilating ducts in the stator
provided for it at the lower part of the stator. Channels are provided under
the floor for both the incoming and the outgoing air.

The desigii of the yentilatixig ducts themselves is a matter upon which much
ingenuity has been expended. The yentilating plate which serves to separate
the stampings, though it should be cheap to manufacture, must be made of such
substantial design that it will not be crushed by the pressure on the punchings, and
Avill not have any parts that can get loose and fly out. At one time ventilating
plates with the spacers punched up were largely in use ; but these are not satisfactory,
unless the metal is thick enough to obviate all risk of the squeezing over of the
punched-up part. Most manufacturers now prefer to rivet spacers of substantial
construction on to an iron punching. There is seldom any advantage in giving
to the internal parts of the spacers the shape of blades in a turbine. The shaping
of the spacers to imitate the blades of a turbine can only be of advantage if the
air is being accelerated in a tangential direction by the spacers themselves. Very
often the air receives its main tangential velocity from the spider arms, and the
shaping of the spacers is in this case of very little use. Most standard machines
have to be designed for rotating in either direction, so that in these it is best
to have the spacers radial (see Fig. 511).

Table XL "Power taken to Drive Fans.

Diameter of

fan in
centimetree.

Outelde
diameter of

frame in
centimetres.

Smallest

opening in

I>ath of air in

sq.cms.

(outlet).

Speed of fan

blade at

1000 B.P.X. in

metres per

second.

Watts taken
to drive fan.

Maximnm

watts carried

away by air fbr

40* C. rise of

machine.

Efficiency

of fanning

action.

37-5

100

250

550

l.OOD

•

13
19
25

40
140
500

2,500

9,500

22,000

12 per cent.

18 M
18 „

214

DYNAMO-ELECTRIC MACHINERY

I
CO

VENTILATION 215

Power takra to drive the fuL The amount of power taken to drive & fan such as
that illustrated in Figs. 215, 218 and 429 depends very largely upon the amount
of air passing through it, and this again depends upon the openingB provided for

pagses from the top of the atator

along coaceatric ducta between the etator punchlngs, and la expelled at the ba«e, aa will be
aeen (rom the longitudinal sectlDa. Scale I: HA,

the air. Id many cases the aii is throttled mainly at one place ; it may be at
the entrance or the exit openings, and often the amount of air is controlled by the
size of these openings. In order to give some idea of the amount of power taken
to drive fans on machines of various sizes, we quote in Table XI. figures for three
frames of difierent sizes.

216

DYNAMO-ELECTRIC MACHINERY

For a given maximum velocity of air, the power taken will vary diiectly as
the amount of air supplied. But for an outlet of given size, the power taken will
vary as the cube of the amount of air supplied per second, because the pressure
required varies as the square of the velocity. For a given number of revolutions
per minute, the power taken will vary as the 3*5th power of the diameter of the
fan, other dimensions remaining constant. For a given fan the way that the power
will vary as the speed ia increased depends upon how far the air is throttled. If
it is completely throttled (on machines having only small openings for the air and

n
te

/3
It
U

9
8

7
0

)

Is ■

! '

iliUiL

-f

/ /

—

100

too 300 400 500 eoo

Revolutions per minute

700

900

Fig. 222. — Approximate values of friction and windage on engine-driven salient-pole

A.C. generators ; 2=30 cms.

a fairly big fan it is almost completely throttled), the power taken varies nearly
as the square of the velocity. Where the passage for the air is free, it varies nearly
as the cube of the speed of the fan.

The following approximate data are sometimes useful. One watt will give a
rise of 1" C. to one gram of air per second. One lb. of air per second requires
453 watts to raise it 1° C. The volume of one pound of air is

273 + 0." 760

12*5 X — z=^^ - X

273 m.ra.

cubic feet,

where m.m. denotes the barometric pressure in millimetres of mercury and
C.° denotes the temperature of the air. If we take the volume of the air

VENTILATION

217

at 35** C. we get the following rule for calculating the amount of air required

for cooling: ^ , . ^ , watts lost

Cubic metres per second = ; ^ — -. ttkk'

'^ temp, rise of air x 1130

One generally allows 50 % more than this in cases where some of the air comes
out at a temperature lower than the maximum.

Friction and windage losses. It may be as well to deal here shortly with
friction and windage losses. It is impossible to compute accurately these losses,
because such small variations in the design often make great differences, especially

100 200 300 ^fOO 500 600 700 SOO 900 WOO ttOO fZOO T300 1400 1500 /600 nOO f800

Revolutions per Minute

Fio. 223. — Approximate values of friction and windage of rotors of induction motors.

in windage losses. Still, as the electrical designer has so often to fill in an approxi-
mate figure for the friction and windage losses in calculating efficiency, it is well
to have a few curves such as those given in Fig. 222 and 223 to aid him. Fig. 222
relates to the friction and windage losses of engine-driven generators. These are
based on tests upon 50-cycle generators having an axial length of 30 cms.

It is foimd that the windage of the rotor of an induction motor as ordinarily
constructed is less than the windage of a salient pole generator of the same
diameter and length. The friction of the bearings is also rather less, because
these are not so massive as generator bearings. Figure 223 gives us rough figures
for the friction and windage of induction motors of standard design. The
curves are marked with the diameter and axial length of the motors.

These curves are only intended to give one a rough idea of the friction and
windage on a machine. The only accurate way of determining these is by actual
measurement.

CHAPTER X.

THE PREDETERMINATION OF TEMPERATURE RISE.

The determination of the temperature rise of any part of an electrical machine
from the design data and a supposed knowledge of the conditions under which
it is worked, will always be a difficult matter ; and no very great accuracy can be
expected from such calculations, because of the impossibility of telling beforehand
exactly what the losses will be, or of predetermining with accuracy the cooling
conditions.

Nevertheless, it is worth while to make a very close study of the ways in which
the heat generated in the iron or copper is carried away, and to make our rules
for the quantitative determination of the amount of heat passed from one part
of the machine to another as accurate as they can be under the circumstances.
Such a study generally leads to a knowledge of defects in the design which can be
remedied. There is no doubt that the great increase in the output per lb. of
material that has been made during the last few years in running machines has
been obtained more by improvements in the methods of cooling than in the reduc-
tion of the losses. The heat produced in any part has a definite path from the
point of origin to the place where it is thrown out from the machine. Thus some
of the PR losses in the armature conductors may have only to pass through a
certain thickness of insulation to the air surrounding the coils; while the heat
generated in the copper in the slots passes through the insulation to the iron, where
it meets with the heat produced in the iron, and both together are conducted to the
ventilating ducts and carried by the air to the exterior.

We can imagine lines of heat flow drawn through the machine which follow
everywhere the paths of the heat from the point of origin to the point of discharge.
At some points there may be constrictions in the path which it is desirable to
avoid ; at others the heat stream flows easily without undue temperature gradient.
Everywhere, at right angles to the lines of heat flow, we can imagine isothermal
surfaces constructed which enclose the points of highest temperature.

In those parts of the machine where there is a heavy temperature gradient,
that is to say, where the isothermal surfaces are crowded together, the designer
must consider what can be done to open these surfaces out, and lower the internal
temperature.

THE PREDETERMINATION OP TEMPERATURE RISE 219

We propose to give in this chapter rules which will enable us to calculate the
amount of heat carried from one part to another under given conditions.

We wiU have to deal with the passage of heat by conduction, by convection,
and by radiation.

Oondaction of heat. It is well to have mental pictures of the relative heat
conductivity of the different materials with which the designer has to deal.

rottc

Heat FUtx

womats

Htat Flux
too Watts

ioo*c

rS7*C

V 77 ins ^'

Fig. 226a.

Figs. 225a, 6, c and i show the heat conductivity of copper, iron, paper and
baffled air.

In these figures the heat-flow is given in watts, that being the most convenient
way of measuring it for our purpose.*

In Fig. 225a we have a copper bar 7 -7 ins. long and of 1 sq. in. section. It is
supposed that the bar is surrounded by a perfect heat insulator (or it may be by
other bars having the same temperature distribution), so that no heat escapes
at the sides. If heat flows in at one end
at the rate of 100 joules per second (i.e.
100 watts), the temperature gradient in
the bar will be as depicted in Fig. 225a.
Two points 7*7' inches apart have a
difference of temperature of 100° C. There
is a difference of temperature of 13
degrees for points 1 inch apart in the
line of the flow of heat. That is to say, a
difference heat potential of 13° C. will
drive 100 watts across an inch cube of j^_. /.7j* ...,*

^^PP®'- FIG. 22».

Fig. 2256 shows a wrought-iron bar of
the same cross-section, along which 100 watts is passing by heat conduction. It
will be seen that the temperature gradient is more than 4 times as steep. In a
bar of cast iron the gradient would be much steeper still. The conductivity of

* The relation between the heat iinitfl is as follows : 1 gram calorie (the heat required to
raise 1 gram of water I'^C.) is equal to 4*2 joules or 4*2 watt seconds, so that the passage of
1 calorie per second through any given surface is equivalent to the passage of 4*2 watt« through
that surface.

HetU Flux
lOOWdtts

fOO Watts ~

220

DYNAMO-ELECTRIC MACHINERY

cast iron depends greatly on the nature of the crystallization. Common cast iron
has only one-half the conductivity of wrought iron.

2SS*C

200*C

r100*C

Heat

Vioo*c

Fl^/

m^ffVatt

HeatFlMX /rp

Fig. 2260.

t^oimut

•5''!<

Fia. 225il.

Table XII. Heat Conductivity of Metals.

Materiau

Copper. (See note on p. 229) -
Steel punchings along laminations
Steel punchings across laminations (10

per cent, paper insulation)
Steel punchings across laminations (8

per cent, paper insulation)
Steel punchings across laminations (7

per cent, varnish insulation) -

Cast iron

Brass

Thbbmal Conductivity.

For square centimetre per * C. of

difference of temperature per centimetre

length of path.

In calories per
second.

0-72 to 10
015

0 0028

00035

00061

003 to 006

0-2

In watts.

3 0 to 4-2
0-63

00118

0015

0026
0125to0-25
0-84

Per square inch

per'C.

of difference of

temperature per

inch length of path.

In watts.

7-6 to 10*6
1-6

0 03

0038

0065
0-32 to 0-64
214

Fig. 225c represents the case where heat flows through pressed paper insula-
tion at the rate of one watt per sq. in. Here the temperature gradient is exceed-
ingly steep, although we are only passing 1 watt per sq. in., instead of 100 watts,
as in the cases depicted of metal bars. If, instead of solid pressed paper, we have
a number of sheets of paper with layers of air in between them, the temperature

THE PREDETERMINATION OP TEMPERATURE RISE 221

gradient will be steeper still. One of the worst heat conductors known is air
which is prevented from circulating by being mixed with some finely divided
fabric.

Fig. 2254 shows the temperature gradient in baffled air. Here we have again had
to reduce the heat flux (this time to 0 1 watt), to make the figure on a reasonable
scale. It will be seen that the heat conductivity of paper is 8^ times as great
as the heat conductivity of baffled air.

Tables XII. and XIII. show the heat conductivity of various materials used
the construction of electric machines.

Table XIII. Hbat Conductivity of Insuulting Materials.*

How Mounted.

Thermal Conductivity.

Materiau

Per square centimetre per

' C. of difloronce of

temperature per

centimetre length of path.

Per square
inchner'C.
of dlfferenoe
of tempera-
ture per
inch length
of path.

In calories
per second.

In watts.

(1)

(2)

(8)

x;'

Varoished oloth (em-

16 turns, each 0*0175 cm. thick.

0 0006

00025

00063

pire doth)

very tightly wrapped

Press-spahn, untreated

2 pieces, each 0-16 cm.

0*00041

0*0017

0 0042

Rope paper, untreated

24 turns, 0014 cm. thick, tightly
wound

000028

00011

0-0029

Rope paper and oil -

24 turns, 0*014 cm. thick, tightly
wound

000034

0*0014

0*0037

Rope paper, treated

Successive turns, 0-019 cm.

000040

00017

0*0042

with sterling varnish

thick, tightly wrapped

FuUerboard, varnished

Successive turns, 0*028 cm. thick,
tightly wound

0*00034

00014

00035

Empire cloth and

Alternate turns of empire oloth.

0*00050

0-0021

00053

mica

0*018 cm. thick; and mica,
0*075 cm. thick, tightly wound

Empire cloth, mica

As in Fig. 227, containing some

000036

0*0015

0*0038

and tape

air spaces

Paper and mica

Allowing for some looseness in
slot

0*00029

00012

0*0031

Pore mica

3 pieces, each about 0*13 cm.
thick

000087

00036

0*0091

Built-up mica -

Bficanite tube containing 19 per
cent, shellac

000025

00010

00026

Built-up mica -

Mioanite tube containing 11 per
cent, shellac

000029

00012

00031

linen tape, treated •

Treated in insulating varnish and
baked

000036

00014

00037

*For method of testing and further particulars see " Heat Paths in Electrical Machinery,"
Symons ft Walker, Jaum. Inst. Eltc, Sngrs,, vol. 48, p. 674.

222 DYNAMO-ELECTRIC MACHINERY

The amount of current that an armature coil or field coil will carry without
exceeding its guaranteed temperature rise greatly depends upon the heat con-
ductivity of the materials with which the coil is insulated, and upon the way in
which they are applied. In many cases it is well to make a rough calculation of
the number of watts per square centimetre which can be passed through the insula-
tion employed under the running conditions. The important factors involved are :

(1) The nature of the insulation and its heat-conducting qualities.

(2) The thickness of the insulation.

(3) The resulting diflference in temperature between the copper and the iron.

Where the insulation is well pressed and in close contact with the copper and
the iron, the temperature gradient within it will be fairly definite and of a known
amoimt, depending on the material. If A^ is the heat conductivity expressed
in the units employed in the second column of Table XIII., the formula connecting
the various quantities is as follows :

Watts per sq. cm. passing from copper to iron

where c is the thickness of insulation in centimetres, 6^ the temperature of the
copper in degrees Centigrade and O^ *^® temperature of the outside of the coil.

Example 29. An armature coil of an induction motor is insulated by means of a tube of
built-up mica, 0*3 cm. in thickness, which fits tightly in the slot. If the permissible running
temperature of the copper is 75° C. and the temperature of the iron is 55** 0. , how many square
centimetres per watt must we allow for the cooling of the coil ?

^i - ^2=20*0. From Column 2, Table XII., we may take Xfc=0*001.

20
Watts per sq. cm. =0'001 x.^.q= 0*067 watt per sq. cm.

For a mixture of paper and mica (half and half by volume), and allowing for
the average amount of looseness which occurs in a well pressed coil, the constant
A^ may be taken at 0*0012 in cm. measure. In inch measure \l is 0*0031.

Example 30. An armature coil is wrapped on the straight part, which lies in the slot, with
paper and mica in equal proportions to a thickness of 0'06^. For a difference of temperature
of 22* C. Iwstween iron and copper, how many watts per sq. in. will pass through this insulation?

Watts per sq. in. =0*0031 ^ iTiH5 = 1 "^ watts per sq. in.

This would be the usual allowance for a direct-current armature coil insulated for 500 volts.

Example 31. A long coil consists of four conductors each 0*6" x 0*3* (say 0*175 sq. in. area).
The wall of insulation consists of 0*125 inch of ptiper and mica in equal parts well pressed and
making a reasonably good fit in the slot. If the copper is worked at 2000 amps, per square
inch, what will be the difference of temperature between iron and copper?

The resistance of the conductors when hot will be about 0*00064 ohm per foot. The current
per conductor will be 350 amps. The total loss per foot run will be

350 X 350 X 0*000054 x 4=26*4 watts.

The mean area of the insulation of a foot n:n will be

9Q*4

4*8" X 12" =57*5 sq. in. ^^jr:p=0*46 watt per sq. in.

0*46=0*0031 x^^\j,^^
Therefore ^i - 6,^= IS'd" C.

THE PREDETERMINATION OF TEMPERATURE RISE 223

In the above examples we have made allowance in the value of k^ for the air
spaces occurring in the insulation. In cases where hard-pressed insulation, such
as press-spahn is employed, there will usually still be a small air space between
this and the adjacent metal, and often this air space is a greater hindrance to the

JOOO -

•oo
eoo

TOO

«oo

SCO

40Q U

»oo

coo

lOO

u..

^
^

*
<•

•

•
-« -

■1/

/
•

•

f
1

/ ,

1 /

1 /
1/

Wid

rhf

Aire

»—

in o<

■nfiiiM

> 1 , 1

»rf>M

o>i

04 o>a 0-4 o-s o«

o-s o«

I'O

I'l

I-*

i-ff

FlO. 226. — ^Thermal reaistaDoes of air spaces of difTerent thicknesses.

passage of heat than the solid insulation. If the amount of air space is known,
or can be approximately guessed, its thermal resistance can be allowed for by
taking values from the curve given in Fig. 226.

Example 32. Suppose that we have a field ooil which is insulated on the inside next the pole
with treated fuUerboard of a thickness of 0*2 cm. From Table XIII. we find that the thermal
conductivity of this material (in watts per square centimetre, etc. ) is 0*0014. The thermal resist-
ance of 1 sq. cm. is there 0*2-^0*0014 = 142, so that if there were no air space and we were
passing to the pole 0*15 watt per square centimetre, the difference in temperature of pole and
coil would be only 21 "S" C. If now we introduce an air space of 1 mm., whose resistance from
Fig. 226 is about 200, the total resistance is raised to 342 and the difference in temperature for
the same heat flow would be 51*5*" C.

Of all the materials used in the insulation of armature and field coils, pure
mica in its original crystalline form is the best heat conductor. If, however, the
mica is split up into laminae, and built up in the form of micanite, the thin layers
of shellac and air enormously increase the thermal resistance, so that built-up
mica is a rather worse heat conductor than many of the fibrous insulating materials.
Indeed any heating or bending of the pure mica, which will interfere with its
solidity, will greatly increase its thermal resistance.

In Table XIII. it will be seen that in several cases (such as the first item) the
thermal conductivity is given for very tightly wrapped material. For the experi-
ments in which the conductivity was measured, the material was specially wrapped
with great care, so as to exclude practically all the air spaces ; so that the values
in these cases must be taken as the maximum obtainable, and must not be used
in practical calculations unless the construction is such as to exclude all air. The
impregnation of coils greatly improves the heat conductivity by filling up air
spaces.

224

DYNAMO-ELECTRIC MACHINERY

Sometimes coils are impregnated with petroleum residue before the main slot
insulation is wound on. This ensures good heat conductivity up to the inside
surface of the insulation, but we have still some air spaces between the layers of
insulation which are wrapped on, and there must necessarily be some little space
here and there between the outside of the insulation and the walls of the slots.
In the machine, the test of which is described below, the insulation was of this
type.

Example 33. A teat was made on a 5000 k.w. three-phase generator by means of thermo-
couples placed in the armature coils during the course of construction. Fig. 227 shows the
arrangement of the armature coils ; the position of the thermo-couples is indicated by the
letters /?, S, T^ Uj V. Junction B gave the temperature of the copper inside the slot ;
S the temperature of the iron surrounding the slot ; 7' the temperature of the outside of
the coil on the part exposed to the air ; U the temperature of the copper in part of a coil
projecting 6 in. from the iron ; V the temperature of the copper in part of a coil projecting 9 in.
from the iron. The generator was run at full speed with the armature short circuited, the
field current being increased until the armature current was 328 amperes. The run was
continued until the temperatures of all parts were constant. The table below gives the degrees,
rise above the temperature of the air admitted to the machine (23" C).

" C. Riae.
i?=390
5^=18-4
r=24-6
tr=38-0
r=.34-4

Fig. 227 gives the arrangement of the conductors and insulation in the slot. It is di-awn
full size. Each conductor, which consisted of two copper straps each 045 in. x0*2 in., was

insulated with tape and mica, a piece of mica 0*03 in. thick
l)eing added as a spacer. All four conductors were impregnated
in vacuo and wound over with empire cloth and mica to a thick-
ness of 0*13 in. The whole was then wound with linen tape.
The total thickness of insulation amounted to 0*177 in. The
various insulating materials were then present in the following
proportions : Empire cloth, 0*07 ; mica, 0*03 ; varnish and air,
0*02; paper, 0*017; tape, 0*04. The heat . conductivity of the
insulation is easily calculated from the above figui-es. The total
loss in the copper conductors per foot run of coil was 27*2 watts.
In calculating this, allowance has been made for the rise in
temperature of the copper and for eddy currents * produced in
the conductors. The difference of temperature between the
copper and iron is 20*6° C. Mean perimeter 5 in., so that the
total area of insulation per foot run is 60 sq. in. With 27*2 watts
per foot run this gives just over 2*2 sq. in. per watt. The specific
conductivity for heat of the insulaticm works out at 0*00153 watt
per centimetre cube per degree. This conductivity is consider-
ably lower than the figure (0*002) found from tests on empire
cloth and mica wound on a copper cylinder with the fewest
possible air spaces, as can be easily understood.

FlQ. 227. — ^Arrangement of
iDBulation in heat conductivity
test.

With coils of rectangular section wrapped with empire
cloth and mica, or paper and mica, in the ordinary
method, one may expect to have a heat conductivity

• See paper by A. B. Field, Jourtial of the American histitiUt of Electrical Engineers^
July, 1905.

THE PREDETKRMINATION OF TEMPERATURE RISE 225

not higher than 00015 watt per cubic centimetre per degree, and for thin insula-
tion, such as used for 500-volt machines, one may take the figuie for A, as

PlQ. £28. — PoalUoni ol thermo-conplea for ttat od tbe htatlng o[ amutiue coJIa,

00012 to allow for the relatively greater importance of looaeneas in the slot.
This in inch measure gives us Aj =00031. For instance,

BxAUPLK 34. On thp armature of a direct -current generelor whose oondiinWirB were
insulated with manilla paper and mica to a thickness uf 0*16 cm., tlie temperature rises
after a full-load run under conditions which mode the squnre inches per wntt 0-9, were
as follows: IiUcmul copper, 41° i iron, 22°. If we use the figure 0-0012 watl per cubic
cenCinietre per degree, we would obtain a turaperature rise of copper above iron of 23°.

Oondoctloii of heat along conductors. It sometimes happens that the copper
conductors on an armature or field-magnet are grouped together so closely that
very little air can circulate between them, and the total cooling surface of the
group is too small to dissipate the heat generated in it- In this case one relies
mainly for coohng upon the conduction of heat along the conductors to parte of
the coils where the cooling conditions are better. A good illustration of this case
is oSered by the end windings of a two-pole field-magnet for a turbo -generator,
such as is shown in Fig. 369. These end windings are completely covered in by n

226

DYNAMO-ELECTRIC MACHINERY

steel end bell, bo that in any case the air would not circulate well between individual
coils, and to avoid the accumulation of dirt it is sometimes found advisable to fill
the interspaces with suitable insulation. A great proportion of the heat generated
in these end windings is conducted along the copper into the parts of the
coils lying in the slots, and from thence it is conducted into the iron of the
field-magnet.

The flow of heat from the centre of the coil to the cooler parts can only occur
if there is a considerable temperature gradient in the end windings. It is necessary
sometimes to calculate what this temperature gradient will be, and what the maxi-
mum temperature rise will be in the centre of the group. The problem is some-
what complicated by the fact that the resistance of copper changes with temperature,
and one ought to take account of this change of resistance because it makes the

FlQ. 229.

watts lost increase according to a compound interest law. Moreover, in most
cases that arise in practice, part of the heat is radiated from the surface of the coils,
and part is conducted along them.

We will first take the case where a conductor is so surrounded by other con-
ductors at the same temperature as itself that the whole of the heat generated in
it is conducted to the cooler ends, and none passes to the sides. Afterwards we
will take the case where a considerable fraction of the heat passes out to the sides
and the remainder along the conductor.

Let M be the centre point of a symmetrically situated end connector so sur-
rounded by other conductors that all the heat generated by electric current in it
passes to the ends. M is supposed to be the hottest point, and from it heat flows
to the right and to the left as indicated by the arrows. It is sufficient to investigate
the distribution of temperature on one side, say the right side. Let the distance
in centimetres of any point P from M be denoted by x. Let the cross-section
of the conductor be 1 sq. cm., so that the volume of any element of length dx is
dx cubic centimetre. Now, as the resistance of copper is almost proportional to

THE PREDETERMINATION OP TEMPERATURE RISE 227

its temperature measured from an artificial zero 240° C. below 0° C, the resistance
of a centimetre cube may be taken to be :

P 16xlO-«x T

^~ 240 '
where T is its temperature in ° C. above the artificial zero.

If /^ is the current density in amperes per square centimetre, the loss per cubic
centimetre will be : rs n t« 16 x IQ-^ x T

The amount of heat passing through the centimetre of cross-section at the
point P will be the sum of all the heat produced between M and P — ^that is to

say- j2 l-6xlO-«f«'^,

Now the heat conductivity • of copper is such that when there is a difference
of temperature of V C. between opposite sides of a centimetre cube, the flow of
heat through the centimetre arrear is equivalent to the ^ ,

heat produced by 3 watts (see Fig. 230). Therefore three • ,

times the temperature gradient gives us the heat flow per I
square centimetre in watts. As x increases the tempera- y^

ture decreases, so that -r- is negative. Thus we have

AT ,a 1-6 X 10-«f*

dx " 240 Jo
We may take as a solution: jr=jrnu,,cos»fl5. _ _

•^ max r yjQ 280.

In cases which we work out in practice the angle "px never assumes values
which make cos^ negative, so that T is always positive. If T were negative
it would be below the artificial zero. The above solution would only be wholly
true if the resistance of copper were negative below this artificial zero.

The distribution of temperature in a conductor such as we have supposed is
therefore given by the top part of a cosine curve, as shown in Fig. 229.

The value of p is

«'<'£S"'-'" -«'"""•■

Therefore 1\ = r^cos(4-7 1 x lO'^ x 7rf x a?),*

where 1^ is the current density in amperes per square centimetre,

X is the distance from the hottest point in centimetres,
T, is the temperature, above - 240** C, at any point «,
r„^x the temperature, above - 240° C, at the hottest point.

An example will make this clearer. Suppose that we have a hot-bed of con-
ductors so bulky that we can assume that the centre conductor parts with no heat
laterally. All heat generated in it passes by conduction to points 20 cms. away

* As the authorities differ as to the heat conductivity of copper, the author has taken a

ue given by Lorenz, which appears to be on the low side. Tests made by the American

Westinghouse Company indicate that the figure 3 '8 watts per sq. cm. per °C. per cm. is more

nearly correct. This would give the formula :

Tx = 7*m*xCOS(4-2 X 10-' X /rf X x).

228 DYNAMO-ELECTRIC MACHINERY

from the centre, which we will suppose are maintained at 10° C. Each conductor
is O'l sq. in. section, and carries a current of 250 amperes. What is the tempera-
ture of the hottest point ?

I^ = 388 amperes per square centimetre.

r.^ {40 -1-240) = 280.
280= 7'„„co8(4-71 X lO"" x 388 x 20).
280 = J'„„cos0-366 = 0-935 r„^.
7",^ = 300.
300 - 2iO = 60° C. ia the temperature of the hottest point.
Now consider the case where part of the heat generated is radiated from the
surface of the group of conductors, and part is conducted to the ends. In the
cases which occur in practice there is a certain specified temperature on the outside

px tn radians
Fia. £81.

of groups of coils which must not be exceeded. Assuming in the first instance
that the temperature is reached, we can roughly estimate the number of watts
per square centimetre which will be dissipated from the surface, having regard
to the thickness of insulation and the amount of air circulation. Let IF represent
the total watts lost in the group of conductors, and w the watts dissipated from
the surface. Then Vi — w will be the heat watts conducted along the copper. The
temperature rise of the hottest point will be lower than if no heat were lost laterally.
Let U3 say that the temperature rise is the same as it would be if the cnirent density
were reduced from /^ to I, and no heat were lost laterally.

From the value of /, thus obtained we can as a first approximation find the
temperature of the hottest point by the foregoing formula, and get a fair idea of
the mean temperature of the whole coohng surface. We can then make a more
accurate estimate of w, and, if necessary, recalculate 7,, and from it T^^.

For convenience in obtaining the values of cos fz from the values of px expressed
in radians, it is well to have a curve such as that plotted in Fig. 231.

THE PREDETERMINATION OF TEMPERATURE RISE 229

Oondttction of heat along poles. It is sometimes useful to make an estimate
of the amount of heat conducted away along an iron pole piece. In most cases
only the roughest estimate can be made of this, because the distribution of tem-
perature is usually too complex for us to get accurate data with which to start
our calculation. If, however, we begin with the assumption that a certain total
number of watts will pass through the internal insulation of the field coil for a certain
average temperature of the pole pieces, we can arrive at a rough estimate of the tem-
perature of the pole surface necessary to dissipate these watts to the air by the
methods considered under the heading " Cooling by air." Having now provision-
ally fixed the average temperatures of the surfaces where the heat is received, and
where it is discharged, it is an easy matter to calculate whether the difference
of temperature is sufficient to drive the heat along the pole. If it is not, then we
must correct our assumption as to the amount of heat coming from the coil, coming
nearer at each trial the average temperature of the inside surface of the insulation
and the radiating surface of the pole.

Cooling by air.*^ There are three main cases occurring in electrical machinery
in which it is necessary to calculate the rate of convection of heat from a solid
surface to the surrounding air.

(1) We have the case of an armature or field-magnet of approximately cylin-
drical shape revolving within the stationary part of the machine. (Cooling
coefficient denoted by h^.)

*2) We have the case of a field coil against which a draught of air is blowing.
(Cooling coefficient denoted by ha*)

(3) We have the case of the iron surface of a ventilating duct, through which
the air is passing at a certain velocity. (Cooling coefficient denoted
by K.)

The laws of cooling of the solid surface are different in the three cases. The
first case (the cooling of the revolving cylinder) is very complicated. A formula
for the close predetermination of temperatures would have to take into accoimt,
not only the square inches per watt and the peripheral speed, but also the length
of the air-gap, the temperature and shape of the surrounding objects, as well as
of the air, the nature of the cooling surface, and the rate at which the air in the
gap is changed by artificial ventilation.

For ordinary direct-current armatures surrounded by ordinary field-magnets
with normal air-gaps, and with no more interchange of air than is naturally pro-
duced by the rotation of the armature, the formula given by Kapp,

^(i+oi«)

gives good practical results. Here 0 is the area of the cylindrical surface, W the
watts to be dissipated, v the peripheral velocity in metres per second, and t^ the
degrees Centigrade rise above the surroimding air.

* For the amount of air required and the various methods of ventilation, see page 206 et eeq.

230 DYNAMO-ELECTRIC MACHINERY

Where we are dealing with a cylindrical cooling surface consisting of iron
punchings only, the coefficient (550 in the above formula) should be given a rather
lower value. Perhaps the formula,*

.,N ^o 333 X watts per sq. cm.

^^ ^" (1+0-lt;)

is as near as any formula can be which does not take account of any other condi-
tions than those embodied in its four terms. The same formula may be applied
for calculating the temperature rise of the internal cylindrical surface of a stator,
V being, as before, the peripheral velocity of the rotor in metres per second.

Example 35. The internal cylindrical surface of the stator of a turbo-generator is 2960
sq. in. , and the number of watts of heat flow communicated to the air by this surface is 11 ,700. If
the peripheral velocity of the rotor is 92 metres per second, find the probable average rise of
temperature of the surface of the stator above the average temperature of the air in the air-gap

11,700 ^«,
^ggg^^-g:^=0-61wattpersq.cm.

_^o 0-61 X 333 ^op
^ =1 + 01x92-^ ^•

An actual test, made on a turbo-generator running under these conditions,
showed a temperature rise of 19° C.

Example 36. The revolving field of an a.c. generator has a diameter of 154 cm. and
a speed of 375 b.p.m. The axial length of the armature iron is 29 cm. What number of watts
can be dissipated from the internal cylindrical surface of the armature for a rise in temperature
35*" C. above the temperature of the air ?

The peripheral speed is 30 metres per second. We have therefore

rt__333 X watts per sq. cm.
(r+6-lx30)~ "
Watts per sq. cm. =0*42.
The total surface is 14,000 sq. cm.
Watts dissipated from surface 14,000x0 '42 =5900 watts.

The cooling of field coils. In predetermining the temperature rise f of field
poils we have two distinct problems : first, to determine the temperatures of the
external and internal surfaces of the coil, and secondly, to find the difiEerence
between the temperature of the hottest point in the coil and temperature of the
surface. In the first problem we have a certain number of watts to dissipate
from a surface of a certain area, and we are concerned with the cooling conditions
on that area. In the second problem we are concerned with the heat-conducting
qualities of the coil itself, and the rate of production of heat per cubic inch, or
per cubic centimetre.

Cooling of the surface of the coil. A surface may be cooled either by air
blowing against it or by the conduction of heat to the body of the pole.

Some designers make their shunt coils to be entirely air cooled. They provide
such large air ducts between the coils and the poles that all heat passes to the air.

• For various forms of similar formulae see the paper quoted above, Joum, Inst. Elec, Engrs.,
vol. 48, p. 674.

THE PREDETERMINATION OF TEMPERATURE RISE 231

Other designers make the coils a tight fit on the poles, and rely upon conduction
of a large portion of the heat generated through the insulation to the body of the
pole, whence it passes to the frame or is dissipated from the pole face. These two
cases require rather different treatment.

In considering the cooling of a surface by means of moving air, we see that
any rules that we may have must necessarily be of very limited application, and
when applied to coils of complicated shapes in proximity to various obstructions,
they can only give us a very rough idea of the temperature that a surface will
attain. Reliable data can only be obtained from experiments on similar coils
run under similar conditions. Still a rough rule is better than no rule at all, even
if it is only of service in indicating the direction along which we may improve the
design. In the matter of air cooling, stationary coils and revolving field coils come
under different rules. With stationary coils we are dependent for our cooling

o^m

ocod

OiW

00^

onee

no. 232.— Belation between hd the watts per sqiure
centimetre per " C, and velocity of air when air
blows upon a cylindrical coil.

Fig. 233. — ^Relation between hd the watts per square
centimetre per ° C, and velocity of air when air
blows upon a cylinder of tarnished brass.

upon the movement of the air, either by the fanning action of the armature or
by some external agency, and the number of watts dissipated per sq. cm. will
depend upon the velocity of the draught against the coils. For coils of approxi-
mately cylindrical shape, which present a surface of cotton-covered wire, the
relation between A^, the watts per sq. cm. per ° C. rise, and the velocity of
the air impinging upon the side of the coil, is given by Fig. 232. The little circles
give the results of a number of tests made on coils with the air blowing on both
sides. The equation ^^ ^ q.qqj ^ ^ ^ Q.g^^j

gives approximately the law. We see that v comes into the equation in the second
power, because, as we increase the velocity, not only do we increase the supply
of air, but we increase the intimacy of contact between the air and the surface.
In the case of draught of air blown in a direction parallel to the cooling surface,
the cooling is proportional to the first power of v. Where the air is blown on only
one side, the cooling effect is greatly dependent on the shape of the surrounding
surfaces. - We may take for the ordinary arrangement of field coils on a continuous
current generator with the air blown from one side,

A^ = 0-001 l(l+0-47t;2).

232 DYNAMO-ELECTRIC MACHINERY

Where the air blows upon a bare metal surface, the cooling is much more effec-
tive. Fig. 233 shows the results of tests upon a tarnished brass cylinder with the
air blown from both sides. The law is approximately

A^ = 0-001 l(l+0-78t;2).

When there are ventilating ducts between the field coil and the pole, the cooling
in the inside surface will be proportional to the velocity of the air in the duct. As it
is impossible in most cases to find out what this velocity will be, the cooling con-
stants can only be determined by experiments on coils of the same type running
under similar conditions. The draught along these ducts is in most cases so low
that the rate of cooling cannot be taken at more than 0*0012 watt per sq. cm. per
° C. rise of temperature. For this reason many designers prefer to do away with
the duct between coil and pole, and cool the inside surface by conduction of heat
into the pole. Even with a thickness of insulation (treated press-spahn) of 0*2
cm. and a liberal allowance for resistance of unavoidable air spaces, we can easily
pass 0*003 watt per sq. cm. per ° C. difference of temperature. With thinner insula-
tion and some care in eliminating the air space, we can pass as much as 0*007 watt
per sq. cm. per degree.

Botating field coils. The cooling conditions with rotating field coils are
usually very much better than with stationary coils ; nevertheless some care must
be taken in the design to take full advantage of the circulation of air set up by the
rotation. Ample space must be allowed for the air to get in between the coils
and any obstructions to free circulation must be removed.

In rough calculations of the heat dissipated from the surface of revolving field
coils, it is usual to take the total surface (both on the exterior and on the inside
next to the pole) and to allow so many sq. ins. per watt. This method is good
enough when we are comparing machines of the same general proportions and
construction, and when we can get frequent check data from machines that have
been tested. The method, though quick and handy in practice, does not help
us to see how the cooling conditions will be altered when the design is modified.

In this rough method of calculation the following figures may be taken for
coils of ordinary construction and well ventilated, with a speed about 5000 feet
per minute, where the length of the poles is about equal to the pitch. For cotton-
covered wire coils an allowance of from 1*2 to 1*4 square inches per watt will give
about 40° C. rise. Where the coils are of bare copper strap, the allowance may
be reduced to 0*8 to 1 sq. in. per watt.

It is better, where time permits, to take separately the cooling of (1) the ends
of the coils exposed to the full draught, (2) the sides of the coils that lie parallel
to the axis of the machine and (3) the internal surface lying next to the pole.

The ends of the coils. The area of the end will be taken to be the area obtained
by multipljdng the length e (Fig. 234) by the height of the coil, and in cases where
the top and bottom of the ends are exposed to the air, their areas should be added.
As a rule, the ends of the coils are flanked with fibre cheeks, which come against a
support. Where this is the case, a short method, of sufficient accuracy, is to take
the heat conducted through the cheeks as equal to the extra heat that would be
conducted to the body of the pole, if the coil were a few centimetres longer than

THE PREDETERMINATION OF TEMPERATURE RISE 233

it actually is. A rough guess can be made as to the number of centimetres to add
to the length of the coil for this allowance as in the example given below.

If we denote by \ the watts per sq. cm. per ° C. rise dissipated by the ends of
a field coil revolving at a radius r^ centimetres, we find that the formula

A, = 0*0011(1 + 1-2 X Rp^ X jRpn, X re X 10"*) watts per sq. cm.

gives us values for the cooling constant which fit very well the results obtained
from tests. In inch measure this becomes

a; = 0-007(1 + 3 X R^^ X sfR^„ x r;' x lO"*) watts per sq. in.,

where r" is measured in inches.

»'^

FlO. 234. — DimensioiiB of a coil upon which its cooling depends ; Id-pole field-magnet.

The rate of cooling of the sides of a coil depends upon the ratio of the distance
« to the length { (see Fig. 235). If we denote by A, the watts per sq. cm. per ° C.
rise dissipated by the sides of the coils, we find that the formula

A,= 1-5 X 10-8 X R^^ X sjR^y, r^ x Jj watts per sq. cm.

gives good practical results. Here r^ is the radius in centimetres. This, in inch
measure, becomes

K; = 3-8 X 10-8 X i2^ X ^^ X r; x .^| watts per sq. cm.

The calculation of the cooling of the internal surface by conduction of heat
to the pole is carried out as indicated on page 223.

The effect of lengthening a frame and of reducing the number of poles can be
seen by the application of these rules to the cases given below.

m one case

234 DYNAMO-ELECTRIC MACHINERY

In Examples 37 and 38 we have revolving field-magnets, each 60' in diameter,
one of 8 inches axial length and the other of 24 inches axial length. The mean
radius of the coil is 25". The clearance between the coils 8 is 0*5 inch. Thus,

yji =0*25, and in the other it is 0'144. It will be seen that, notwith-
standing the much larger cooling surface exposed in the longer machine, the total
watts dissipated per coil are only 778, as against 516 in the case of the shorter
machine.

Example 37. Fig. 234 gives the dimensions of a 16-pole field-magnet. The axial length
of coil is 8'^ The speed of the field-magnet is 375 r.p.m. Find the number of watts dissipated
for 40° C. rise above the temperature of the air. Take the temperature of the pole at 8* C.
above the air. We may take e, the length of the exposed end, at KT.

A«=-0011(l + l-2xiJp«XN/^^xrcXlO-»)
= -0011 ( 1 + 1 -2 X 375 X n/376 x 63 5 x 10-»)
= -0011 (1-f 5-5)

= -0011 X 6-5= -0071 watt per sq. cm. per V C. rise.
Area of ends of coils =2 x 7" x 10" x 2'64* cms. =905 sq. cms.
Watts dissipated at the ends, per coil = 0071 x 905 x 40=256 watts.

A=l-5xl0-«xiJp«XN/^xreX>y'!

= 1-5 X 10-« X 375 X n/375 X 63*5 X J
= -OOnS watt per sq. cm. per 1" C. rise.
Area of sides of coils =2 x T' x 8" x 2-54* x 723 sq. cms.
Watts dissipated at the sides, per ooil= -00173 x 723 x 40=60 watts.

In calculating the heat conducted to the core we may add 1} inches to the length of the
coil to allow for the heat conducted from the ends. As the insulation between coil and pole
usually contains air spaces of uncertain dimensions, we may take the thickness as being
0*25 cm. and the heat conductivity at "001 watt per sq. cm.

w *.* 32 X 001 ,^

Watts per sq. cm. = — ^ — = -128.

Watts conducted to core, per coil = 30 x 2-542 ^ (7 + ij) x '128

=210 watts.
Total watts per coil = 256 -f 50 +210

=516 watts.

Total heat that can be dissipated from field coils = 16 x 516

= 8-3 K.w.

Example 38. Dimensions as in Fig. 234, but ^=24^

Watts dissipated from ends of coils (as above) = 256 watts per coil.

h =l-5x 10-8x37W375x63-5x 144
= *001 watt per sq. cm. per T C. rise.
Area of sides of coil =2 x 24 x 7 x 2 "54* =2170 sq. cms.
Watts dissipated at the sides, per coil= -001 x 2170 x 40=87 watts.
Heat conducted to core, watts per coil =62 x (7+ IJ) x 2-54" x -128

=435 watts.
Total watts per coil = 256 + 87 + 435

= 778 watts.
Total heat that can be dissipated from field coils = 16 x 778

= 12*4 K.w.

THE PREDETERMINATION OF TEMPERATURE RISE 235

Example 39. Same diameter as before, bat 8 poles iDstead of 16 (see Fig 235), speed
375 R.P.M.

,=24" = 61 cms. ^j-l=^^[^='^'^^'

K= -0011 (1 + 1-2 X 375 X \/375 x lO"' x 61)
= •0011(1+5-3)
= 0011x6-3
= •00693 watt per sq. om. per 1** C. rise.

Area of ends of coil =2 x 18 x 8 x 2-54'= 1860 sq. cms.
Watts dissipated at the ends, per coil= 00693 x 1860 x 40=515 watts.

^, = 1 5 X 10-* X 375 X V375 x 61 x -577
= '00383 watt per sq. cm. per 1^* C. rise.

Area of sides of coil =*2 x 12 x 8 x 2*54'= 1240 sq. cms.

Watts dissipated at the sides, per coil = 00383 x 1240 x 40

= 190 watts.

Fio. 235. — Dimensions of a coil upon which its cooling depends ; 8-pole field-magnet.

In calculating the heat conducted to core, we may add 2 inches to the length of the core to
allow for the heat conducted from the ends.

Watts per coil =50x2 542 x (8 + 2) ^ .128
= 414 watts.

Total watts per coil =515+ 190 + 414

= 1119 watts.

Total heat that can be dissipated by field coils =8 x 1119

= 895k.w.

In these examples the temperature of the surface of the coils has been taken at 40° C. above
the air. If this were the case the mean temperature of the coil would be a few degrees higher.

236

DYNAMO-ELECTRIC MACHINERY

Having ascertained the approximate rise of temperature of the surface of the
coil, the next step is to aee how much higher the temperature of the inside of the
coil is. In general, no calculation need he made of this, because the designer
knows from experience of similar cases that the temperature is not too high. How-
ever, if a particularly deep coil is to be made, or one which contains an exceptionally
large number of layers of fine wire, and in all cases where field coils are mn at
temperatures near the danger point, a calculation should be made of the rise of
temperature of the interior of the coil over the surface.

The condnction of heat acioss the layers of insnlated wires ia a coll. The
internal layers of a shunt coil, when heated up by the current, are hotter than
the external layers. In a large number of measurements made by Mr. Rayner ♦
on coils under running conditions, it was found that the temperature reached
by the inside layers was frequently 50° C. higher than the temperature recorded

180'C

lao'c
uo*c
KO'Q

9 L»ngltudin*r Saction, So

Lis i FuU Slja. TV'Mav

flQ. £36.

by a thermometer placed on the outside of the coil. In many cases the maximum
temperature was 20° C. higher than the mean temperature measured by the resist-
ance method.

As it is the maximum temperature reached that determines the Ufe of the coil,
it is important to design the coil so that this maximum temperature will not be
too high. It is also important for the designer to know approximately how much
higher the temperature of the coil will be in its hottest part than on the outside,
so that he may (if he so desire) work the copper at its highest safe current density.

A genera] study of the distribution of temperature inside a coil shows that
the hottest part is commonly midway between the top and bottom of the coil at
a point a little way removed from the iron core. Fig. 238, taken from Mr. Rayner's
paper, shows typical curves of temperature distribution inside a coil operating
under practical conditions. The dimensions of the coil are given in Fig. 237. The
temperatures were measured by thermo-couples inside the coil at the points indicated
by the black dots in Fig. 237. This coil, one of six in a 94 h.p. motor, consisted of
2584 turns of wire 0'075" in diameter double-cotton covered. The whole coil was
impregnated with insulating gum, and covered with tape. ^Vhen curve A was
obtained, the current density in the wire was 980 amperes per sq. in. The total
'Jmr. laxe. Eltc. Bngrt., voL 3*, p. 828.

THE PREDETERMINATION OF TEMPERATURE RISE 237

watts converted into heat in the coil were 407. The machine was then running
at full load at 325 R.P.M., the diameter of the armature being 31*1 inches. As
the total surface of the coil (including the part next to the pole) was 816 sq. in.,
the watts per sq. inch were 0*5. It will be seen that the highest temperature
reached was at a point about V from the core, so that two-thirds of the heat
travelled to the outside of the coil, and the other third (with the exception of some
that came out at the ends) went towards the pole. In any coil the fraction of

ELEVATION

PLAN

Fig. 237.

the heat that goes towards the pole will depend upon the heat conductivity of the
insulation between the coil and the pole and the temperature of the pole. In
this case the rate at which heat was being conducted through the outer layers
of wire at a point halfway between the top and bottom of the coil was about 0*525
watt per square inch. It will be seen from curve A that the temperature gradient
at this point was 40** C. per inch.

Now this temperature gradient depends not only upon the watts per sq. in.
of heat flux across the coil, but also upon the size and shape of the wire, the way
that it is bedded and the nature of the insulation. If we are given full particulars
of the size of wire, the thickness of the insulation, the space factor, the number

238 DYNAMO-ELECTRIC MACHINERY

of turns and layers, the exciting current, and so on, we should be able to predeter-
mine with a sufficient degree of accuracy the temperature of the hottest part of
the coil.

The problem is somewhat analogous to the case already considered where the
heat is conducted along copper conductors, but in this case the heat is conducted
across one layer of conductors to another. The law of distribution of temperature
takes the same general form :

where T^^ is the temperature of the hottest point measured from the artificial
zero (240'' below 0° C), and T, is the temperature of any point distant x centi-
metres from the hottest point along a line drawn in the direction of the flow of
heat at right angles to the cooling surface. The value of p^x in practice is such
that cos pjX never assumes negative values.

If we examine the various curves given by Mr. Rayner * we will see that they
are all part of cosine curves, except in those cases where there is a discontinuity
in the coil.

Take, for instance, Test No. 2b. Add 240° C. to the ordinates of the trans-
verse section curve on page 639 (vol. 34), and we obtain a curve like that given
in Pig. 238. The law of this curve is approximately :

T,=356cos(0-0975rc).

If the coefficient {p^)oix is known,f and the distance from the hottest part is known,
then we can calculate the amount that the temperature of the hottest part exceeds
that of the surface. For instance, with the above law, if on the surface of the
winding the temperature is 90° C. (330° above the artificial zero) and the hottest
point is 4 cm. from the surface, then

330 = r,„»,co8 0-0975 x 4,

^m« = 356.

The value of p^ depends mainly on four factors :

(1) The current density in the copper.

(2) The thickness of the insulation per centimetre depth of coil and its nature.

(3) The space factor of the winding.,

(4) The ratio of the length of the bobbin to the depth of the windings.

* Jotimai of the, Instiiuiion of Ehctriccd Ehgineera^ vol. 34, p. 613. See also G. A. Lister,
** The Heating Coefficient of Magnet Coils," ibid, vol. 38, p. 399.

fit is not possible to predetermine the value of pi in all the oases given by Mr. Rayner,
because full particulars are not given of the thickness of the cotton coverings, but in several
oases where we may assume the cotton covering is normal and the wires properly packed, the
results agree closely M'ith the values of p found by the author's experiments ; for instance, in
the case of coil No. 2 we have

p, = 127 Jg--^^^ ^'^^ '^'^^^-^^ =0-0975.
^^ \ 0-00095 X 240

The value 127 amperes per square centimetre is obtained from the value 151 given in Mr.
Rayner's table by the formuUv for /, given above. Length of coil = 7 in., breadth =3 in. ;
7 + 3 = 10.

^ = 0-84; 151x0-84=127.

THE PREDETERMINATION OF TEMPERATURE RISE 239

In what follows we shall employ the following symbols :

I = length of bobbin in centimetres.

d = depth of winding in centimetres.

la = current density in amperes per square centimetre.

-u

o- = copper space factor.

i„ = thickness of insulation per centimetre of depth of winding.
ks =heat conductivity of insulation in watts per square centimetre per °C,
per centimetre of path.

Then

^i = W-

6 X lb"* X o- X i„

866 r

880,

h X 240

In order to ascertain the values of kf^ for round and for square wire, treated
and untreated, experiments were made on the heat conductivity of cotton-covered
wire windings. In Table XTV. are given values for
k,^ in some typical cases.

The figures given in this table allow a certain
margin for variations in the construction of the
coil which, so far as the tests went, appeared to
be sufficient for tightly wound coils. For instance,
the lowest value obtained for 0*032 in. round wire
double-cotton covered and enamelled was 000065,
and the highest value for 0'114 in. wire was 0'0009.
For untreated wires both sizes averaged about
0*00055. It is possible that the margin given should
be made wider. For loosely wound coils it will be very wide. The value of k^ is
independent of the thickness of the insulation on the wire. The thickness of the
insulation is taken into account in the formula in the quantity i„, which is obtained
by multiplying the number of layers per centimetre with the double thickness of
cotton covering on each wire.

Tabls XIV. Value or kh fob Wirr- wound Coils.

Fig. 288. — Curve ahowing distribution
of temperature inside a shunt ooil.

Kind of wire.

How treated.

Diameter of wire.

Inches.

Square wire double

Made solid with heat-con-

0114

000120 to 000140

ootton covered

ducting enamel

Square wire double

Untreated

0114

000090 to 000100

cotton covered

Round wire double

Impregnated and made into

003 to 0114

0 00086 to 0 00095

cotton covered

solid block

Round wire double

Treated with enamel -

003 to 0114

000066 to 000090

cotton covered

Round wire double

Untreated, tightly wound -

007to0114

0 00050 to 0 00060

cotton covered

Round wire double

Untreated, tightly wound -

003 to 0070

0 00040 to 0 00060

cotton covered

Round wire double

Untreated, loosely wound -

003 to 0070

000020 to 000035

cotton covered

240

DYNAMO-ELECTRIC MACHINERY

Example 40. A shunt coil of a c.c. generator is wound with 3480 tui-ns of round double
cotton -covered wire, dia. 0*080^ bare, 0'09*2" insulated. The dimensions of the coil are as given in
Fig. 239. There are 40 layers of 87 turns each. Between the coil and the pole there is a total
thickness of ^ inch treated fuUerboard, and not more than -^ inch of air space. There is a
fan on the armature which creates a breeze, which is directed by the frame in an axial direction
at a velocity of 2 metres per second against the sides of the coil. The ends of the shunt coil
are flanked with J* press-spahn, and are disposed in such a way that the cooling of the ends may
be taken as about half as good as the sides.

t-3-€

30*

Z€*

5V-*

73'

Fig. 239. — Dimensions of large shunt coil from which the temperature rise inside the coil

can be approximately determined.

Find the maximum temperature rise inside the shunt coil after a long run at 4*35 amperes
exciting current.

The total length of wire in the coil will be about 20,600 feet, having a resistance of 33 ohms
cold or say 40 ohms hot. The total watts lost in the coil will therefore be about 760.

The thermal resistance of the fuUerboard, 0*254 cm. thick, is 180 and of the air space 150,
giving a total of 330 ; so that we have 0*003 watt conducted per sq. cm. per ° C. Now calculate
the cooling coefficient of the external surface. This is

hji =0*0011 (1+0*47x2x2)00032.

\i} ^ for the ends is half this we may conveniently take half the area at 0*0032.

THE PREDETERMINATION OP TEMPERATURE RISE 241

The cooling constants being approximately the same, we will as a first trial apportion the
watts between the surfaces in proportion to their area. The area of the various surfaces are :

Sq. cm. Watts taken away.

Inside surface touching pole 2900 258

Outside surface - - - - - - - - 4000 357

One end surface 1620 145

8520 760

Now find the temperature drop through the insulation with this provisional apportionment
of the total watts :

,-T^Tr^= 0*089 watt per sq. em. ^r^^r^= about 30" C.

If the pole were 35" C. (10" hotter than the air), this would make the inside of the coil next
to the insulation 65" C. Next find the drop of temperature between outside of coil and air :

4000 ~"^*^^* 0 0032-^^ ^•

If the air blown on the coil be taken at 30" C. , this would give 58" C for the running tempera-
ture of the exterior of the coil. Now see if this distribution of temperature will fit sufficiently
well a temperature gradient curve with its apex in a suitable position to give the assumed flow
of heat inwards and outwards. The copper space factor is 0*615. The total thickness of
cotton covering per cm. is 0* 135 cm. The value of kk can be taken from Table XIV. to be 0*00095
The current density 134 amps, per sq. cm. must be multiplied by the coefficient

"8~

0*83

= V8+^
to allow for the cooling towards the ends of the coils. Thus we have

P, =0135 X 0'83>v/Q'^'-^ ^^ -^ ^^^^^'^'^'^^=0*0855.
' \ 0*00095x240 ^^ooo.

The law of the temperature gradient of curve is

T, = 7\a»xCos (0-0855 x a-, ),

where x^ is the distance of the apex of the curve from the outside surface. This distance x^
must be found by trial and error. In fixing provisionally the position of the apex of the
temperature gradient curve, we must remember that it is the watt-shed of the coil. It marks
the position of the surface inside which all heat travels inwards, and outside which all heat
travels outwards. The total volume of the coil should therefore be divided by the watt-shed
plane into two volumes, one of which supplies the heat travelling to the inside and the other
the heat travelling to the outside. If now, in our example, we put the watt-shed surface at a
distance of 4*8 cms., from the outside, we will find that the amounts of heat generated in the
volumes cut off are about in proportion to 357 and 258 respectively. We find Tm^x from

Tx, = 7 m»x cos (0*0855 x 4*8),
where T,, = (58" -f 240") =298". This gives us 7'„« = 325".

Thus the law of distribution of temperature within the coil becomes

r»^(85-f 240) cos (0-0855a:).

From this we find that the temperature of the copper next to the internal insulation works
out to 63" C.

This is sufficiently near the assumed value 65 for us to accept the position taken
for the apex of the curve. Fig. 239 then gives approximately the distribution of
temperature under the prescribed conditions.

The passage of heat from the surfEU^e of ventilating ducts to the air flowing
through. When air is blown through a ventilating duct, the distribution of the
stream lines is usually very complicated. Generally we may say that the air in
close proximity to the walls of the duct has a velocity much lower than the mean
velocity, and the air in the centre of the duct has a velocity above the mean. In

W.M. Q

242 DYNAMO-ELECTRIC MACHINERY

what follows, when we speak of the velocity t; of air in a duct, we mean the average
vehfCity, that in to say, the number of cubic metres passing per second divided
by the area of the cross-section of the duct in sq. metres. In turbo-generators,
and other machines which are fed with a known quantity of air per second, the
average velocity of the air in the ducts can be approximately calculated. But
in ojK!n machines it can only be guessed at. Even if the guess is wide of the mark,
it may be useful in comparing the performance of different machines, so long as
we make the guess according to a definite rule. For instance, if we say that the
Vf;locity of the air in the ducts of the stator of an unenclosed generator or motor
can be taken at one-tenth of the peripheral velocity of the rotor, we have a figure
which, though very far wrong in some cases, nevertheless enables us to compare
the cooling effects of ducts of the normal size in widely different machines in a
more inteUigent manner than if we allow the same specific cooling coefficient for
all ducts, whatever the peripheral speed. The formula given below for the specific
cooling of the walls of ventilating ducts is intended for use in turbo-generators
and other machines in which the velocity in the ducts is approximately known.
We may, however, use it in default of any other for calculating approximate figures
for open -type machines.

From tests * on a turbo-generator, it was found that

Where ho is the watts per square centimetre of cooling surface per 'C, the
difference of temperature between surface and air, v is the mean velocity of
the air in the duct and metres per second, and Kr, is a coefficient.

The value of the coefficient Kv was '0014 in a number of tests in which v
varied widely, and as the formula, applied in the way that we propose, was found
to 1(1 ve the temperature rise with fair accuracy under practical working conditions,
it is of more value than a formula based on laboratory experiments.

The coefficient 0(X)14 is higher than would be obtained if the air were blown
through a flat ventilating duct so steadily as to undergo no disturbance of the
stream lines. Tests made under these conditions with a ventilating duct J" wide
giivi* a coefficient of 0'()0()5. The violent eddies which occur in air when it is
(liMchurut'd from the air-gap into the ventilating duct increase the cooling
pr<»|)eriioH.

A round ventilating hole 2" diameter, through which the air passes with
Htuady stream lines, will give us A,. = "OOOSSv, but if baffles are added to stir up
Uio air tho formula may be A,, -^•()011v.

In applying this formula, v is taken as an average figure for the whole of the
(luetN. In finding the area of the cooling surface, the area of both walls must be
eounled,

Uning the formula in this way, one will, of course, only get the average rise
in temp<»rature of the surface of the iron over the average temperature of the air
in the ducts. Unless care is taken to distribute the cool air evenly, points may
he fotmd whore the temperature rise is considerably above the mean. A good
iilort of the way in which the heat distributes itself in a turbo-generator, and of

• Jouni, IiiHt. Klec, Entfinerrs, vol. 48, p. 703.

THE PREDETERMINATION OF TEMPERATURE RISE 243

the factors which control the temperature rise, is obtained from the following
description of experiments on an 1875 K.v.A. generator running at 3000 R.P.M.
The machine was totally enclosed and ventilated by means of a fan at each end,
in the manner shown in Fig. 215.

In certain parts of the machine ordinarily inaccessible to thermometers, thermo-
couples were placed while the machine was in course of construction. Thus, in
the centre of the packet of punchings lying between the ventilating ducts Nos. 5
and 6 (see Fig. 240), thermo-couples were placed at the points M, N, 0 and P. Then,
in the ventilating ducts at the lower part of the machine, couples were exposed
to the full blast of air, so that readings could be taken of the temperature of the
air in the ducts at that part to compare with the temperature of the air in the
same ducts at the top of the machine.

Fio. 240.— Showing depths H, I, J, K and L, to which thermo-couples were inserted into yent ducts.

The generator was run at no load, and the ironloss, friction and windage measured
in the ordinary way by measuring the power supplied to the driving motor. The
rotor was mounted in its own bearings, and coupled to a driving pulley mounted
on independent bearings. The pulley was driven by a direct-current motor at
3000 revs, per minute. The power taken to drive the pulley alone with the full
tension on the belt was found to be 13 K.w. The sum of the friction of the genera-
tor bearings and the windage was found by deducting 13 K.w. from the whole
combination. With full aperture allowed to the fan, the sum of the friction and
windage amounted to 46 k.w. In order to ascertain the amount of power lost
in the bearings, measurements were made of the quantity of oil supplied to each
bearing per minute and the rise in temperature of the oil. A rough estimate was
also made of the quantity of heat lost by the bearings by radiation and convection.
It was found that the heat carried away by the oil in one bearing was equivalent
to 79 K.W., and from the other 6*7 k.w. The radiation and convection losses of
the two bearings together was less than 1 k.w., so the total bearing losses were
about 15*6 k.w. This left 46 - 15*6 =30*4 K.w. for the windage with full aperture,
giving 8800 cub. ft. of air per minute at 50® C. With a reduced aperture giving
4400 cub. ft. of air per minute, the windage loss was 22*8.

244 DYNAMO-ELECTRIC MACHINERY

The amount of air passed. through the machine per minute was measured in
two different ways : (1) An anemometer was used to find the mean velocity of air
at the exit in feet per minute, and this multiplied by the area of the exit in square
feet gave roughly the cubic feet per minute. (2) The total rise in temperature of
the air in passing through the machine was measured, and from the known losses
causing the heating, the flow of air could be calculated. The first method was
not as accurate as the second. It gave on the average an air velocity from 5 to 7
per cent, too high. We have therefore adopted the figures given by the second
method. These are probably right within 5 per cent. It must be remembered
that what we are really concerned with is the weight of air passed through the
machine per minute. The volume of the air changes with the temperature quite
appreciably. Thus, at 20° C, 750 lbs. of air have a volume of 10,000 cub. ft.,
while at 60° C. the volume is 11,400 cub. ft. There were three tests, which we
distinguish by the letters A, B and C

In test A the air supply was cut down to about half its normal flow. The
field-magnet was excited with 133 amperes (about 30 per cent, more than the
no-load field current). The resulting iron loss was 43*5 K.W., and the l^R loss
in the field-magnet was 8*5 K.w. Thus the total losses going to warm up the
air were :

Kilowatts.

Windage 22*8

Excitation 8*5

Iron loss 43*5

Total - - - - 74-8

After running 4 hours, the temperature of all the parts of the machine rose
within half degree of their final temperature. The air entered the machine at an
average temperature of 21*7° C, and was expelled at an average temperature of
53*2° C, giving a temperature rise of 31*5°. The heated air did not represent the
whole of the heat produced. The cast-iron frame presented a cooling surface of
10,900 sq. in., and had a mean temperature over the air of 28° C, so that it would
radiate * in almost still air about :

10,900 X 0-008 X 28 = 2'44 k.w.

The cast-iron blocks upon which the frame rested would carry away not more
than 1 -5 k.w. Let us say that 4 K.w. was lost by the frame. This is such a small
fraction of the whole that we need not estimate it very accurately. Then we
have 74*8 -4 =70*8 K.w. carried away by the air. Now 1 k.w. is equal to 240
calories per second, so we have

70*8 X 240 = 17,000 calories per second.

* It is of interest to note that a small fraction of the total losses on a large turbo generator
are dissipated by external radiation ; in this case about 5^ per cent. In the case of a medium-
speed generator of 5 K.w. about 50 per cent, of the losses can be accounted for in external
radiation. Radiation is used here in its commonly accepted inaccurate sense, and includes con-
vection to nearly still air. The true radiation by heat waves is rather less than half of these
figures, and may be calculated approximately from the formula :

/Tin gram calories per sec. = surface in sq. cm. of equivalent sphere x(rx (Tj-T^), where
<r=0*6x 10~" for a dark painted generator and Ti = temperature of generator in °C. above
al>solute zero. Ta= temperature of surrounding objects above absolute zenj.

THE PREDETERMINATION OF TEMPERATURE RISE 245

Taking the specific heat of air at 0'2375, we have

17,000 1 1 60 „^^„ - . . .

0-23^ ^ STS ^ 453 ^ T =^ ^^^ ^^'- ^^ ^"^ P^' °"^^*^'

or 4400 cub. ft. of air at 53° C. The anemometer measured on the average 4800
cub. ft. of air per minute. This reading must be too high, because 4800 cub. ft,
of air per minute raised in temperature 31*5° represents more power than was
actually supplied to the machine, so we will take 4400 cub. ft. as about right.

It is interesting now to see exactly how the air was heated up as it passed
through the machine.

The temperature of the air in the various ventilating ducts and in the air-gap
was measured by a pair of thermo-couples, mounted on a long wooden rod, which
could be moved about in the ducts while the machine was running. Two couples
of equal resistance connected in parallel were used, one on each side of the rod,
so that if there were any difference between the temperature of the air on one side
of the duct and the other, the reading obtained gave the average value. The
couples were of such a very thin wire (0*01 in. diameter), and were mounted in
such a way that, when exposed to a breeze, they assumed the temperature of the
air almost immediately. It was therefore possible to take very rapidly a large
number of readings of the temperatures at different depths in each air-duct, and
to plot curves such as those given in Fig. 241. The lines marked H, /, J, K and L
are drawn through the points which give the readings of temperature rise at dif-
ferent depths in the ventilating ducts as indicated by the dotted lines in Fig. 240
bearing the corresponding letters. The hole at the top of the frame at which
the air was expelled measured 36 x 20 in., and it was only over this area that it
was possible to insert the wooden rod carrying the thermo-couples. In some parts
within reach of this hole a flexible strip of press-spahn with a thermo-couple attached
was used to take check readings, and the couples placed in the ducts in the lower
part of the machine (that is to say, below the rotor) were used as a further check.
These lower couples gave readings 2° or 3° higher than couples placed in the same
ducts in corresponding positions at the top of the machine. This was possibly
on account of the slightly lower velocity of the air thrown downwards, there being
a certain amount of back pressure produced by the resistance of the flow of air
through the annular space in the frame. As far as could be ascertained by a
number of check readings taken over the field available from the exit hole at the
top, the chart in Fig. 241 represents fairly well the distribution of temperature in
the ducts in the top half of the machine, and if similar charts had been taken in
radial planes at various angles all round the machine, the chart would have been
very similar, but all the temperatures would have been gradually raised about 3°
as we approached the planes lying below the rotor.

Temperatures were at the same time taken of the air admitted, of the air in the
end bells, in the gap, in the yoke, and at eight different points distributed over
the exit.

The average temperature of the air drawn into the machine was 21*7° C. It
will be convenient to speak of the temperature rise over this initial figure, rather
than of the actual temperature of the air. In the end bell at the points F and G

246

DYNAMO-ELECTRIC MACHINERY

ffO

the temperature had risen 9*8® C. and 10*2° C. respectively. This rise was due
partly to the work done to the air due to the centrifugal blowers, partly windage
and I^R losses from the end bells of the rotor, and partly from heat radiated from

6 7 a 9 iO M

N9 of ¥entihJb/n§ duct.

no. 241.— Temperature rise of air at different depths in the ventilating ducts. Iron loss, 43*5 K.W.

Air supply, 4400 cub. ft. per minute.

end plates of the machine. The lowest temperature recorded in the air-gap was
at the entrance to the first ventilating duct. Here the rise was 14° C. As we
passed from the first vent duct to the centre of the machine the temperature was

e 7 a 9 to
MS of ^nti/ating' oba^*

Fio. 242.— Temperature rise of surface of iron in ventilating ducts.

higher, but the increase followed an irregular law, as indicated by the wavy line
marked position H in Fig. 241. The curious dip in the curve in the centre of the
machine, which is also seen in the curves of temperature rise in the iron in Fig. 242,
only occurred when the air supply was throttled. It does not occur in Figs. 246,

THE PREDETERMINATION OF TEMPERATURE RISE 247

246, and 249. The throttling of the air supply reduced the pressure in the end
bells from 425 in. of water for full aperture to 0*75 in. of water. Thus the blast
along the air-gap must have been very much reduced while the blowing action of
the vent ducts on the rotor would have taken a more important part in the scheme
of ventilation than when the full blast was in operation. Owing to the meeting
of the two opposing currents of air in the axial holes in the rotor, there is a ten-
dency for the pressure of air from the rotor to be greatest in the middle, and this
increased pressure probably gave a supply of rather cooler air near the centre of
the machyie.

The velocity of air at the exits of the different vent ducts, though not perfectly
constant as one passed from duct to duct, was only very slightly greater in the
centre of the machine than at the ends. It is therefore sufficient for our purpose
to take the mean temperature rise of the air entering the ducts as derived from
curve H at 20*5^ C.

Let us now calculate the number of kilowatts, a?, required to heat up the air
to 20*5® C. We have, from our previous calculations,

X 20-5
70-8""31-5'
a; = 46 K.W.

Now the windage only amounted to 22*8 k.w. and the 7*22 in the field to 8'5,
80 that we have 14*7 k.w. in addition which must have been supplied by the iron
loss, and communicated to the air mainly on the cylindrical face of the armature.
A small amount — probably about 3 k.w.* — would be supplied to the air from the
end plates of the armature. Deducting this, we have about Wl k.w. conveyed
to the air by the cylindrical face of the armature. As we have seen above, we
are able from this data to calculate the specific rate of cooling per square inch of
armature face.

As the air passes along the vent ducts, the temperature rises ; in some ducts

the air received as much as 11 '5° further rise in temperature, in others not more

than 8** rise, the mean being about 102° rise. If y is the power expended in heating

up the air 10*2**, we have

y ^10-2

70-8~31-5'

y = 23 K.W.

Now the air passes into the annular space in the frame and picks up a little
more heat from the pimchings. Part of this extra heat is communicated to the
frame and is radiated from the outside, and part goes to raise the temperature to
53*2°, giving a total temperature rise of 31*5°.

The next point of interest is the distribution of temperature in the iron punch-
ings. The thermo-couples were placed in the centre of a packet of punchings at
the points M, N, 0, P (see Fig. 240), another couple was placed at Qy just behind
the first punching in the packet ;• the packet in question was the one between
ducts 5 and 6 in Pig. 240. For the purpose of taking rapid readings of the tempera-

* That this amount is small can easily be seen when we come to calculate the amount of heat
given to the air in one ventilating duct.

248

DYNAMO-ELECTRIC MACHINERY

ture on the surface of the iron punchings within the ducts, an instrument was
made, which consisted of a piece of copper foil 0125 x 0*75 x 001 in. soldered to
a thermo-couple mounted on a velvet cushion, and arranged on a wooden rod, so
that it could be pushed down the ventilating ducts and pressed against the sheet
iron. A spring was provided at the back of the cushion to give the requisite pres-
sure, the copper foil being shielded from draughts by the cushion, and being of
small heat capacity very soon assumed the temperature of the iron against which
it was pressed ; thus one could read off directly on a millivoltmeter the tempera-
ture of any surface against which the copper foil was placed. .

Fio. 243. — Temperature rise inside one packet of Iron punchings 4*6 cm. thick and 29 cm. deep.

Lobs about 0*055 watt per cubic centimetre.

Fig. 242 gives the distribution of temperature of the iron on the surface of the
various ducts in test (a). The curves H, /, J, K and L correspond with the positions
in the ducts shown by the lines in Fig. 240.

It will be seen that these curves follow the general shape of the curves giving
the rise of temperature of the air, but they are, on the whole, about 10*5° higher
for the position H and 8*5° for the position L.

If we take the average value of the temperature of the air at the position H^
then the average value at the position /, and so on, and plot these average values,
we get a curve giving the mean temperature rise of the air as it passes through
the ventilating ducts, like that shown in Fig. 244. The ordinates in this figure
give the rise above the temperature at which the air enters the machine, the rise
before reaching the ducts being 20*5°, and the rise in the ducts being 10*2° C.

Taking similarly average values of the temperature of the iron at the various
positions, we get a curve of temperature rise of the surface of the iron.

THE PREDETERMINATION OF TEMPERATURE RISE 249

As the total amount of air passing through the ventilating ducts was 300 lbs.
per minute, it is possible to calculate the watts absorbed by the air from the rise
in temperature by the formula :

rri ^x 300 X 453 X 0*2375 x temperature rise
^''°**"' = 60x240

Plotting the kilowatts absorbed by the air, we get the curve shown in Fig. 244.
The velocity of the air in that part of the ventilating duct which was narrowed by
the armature coils was 8*4 metres per second, in the part of the ventilating duct
beyond the armature coils the velocity was 4*5 metres per second, and at the exit
of the ventilating ducts the velocity fell to 3*2 metres per second. Plotting these
figures, we get the velocity curve in Fig. 244. We have given, on the same figure,

^^00 cuJb/c feet ofaM'/Der minute

Dfatence from A/r gvtp /n cen^metrea.

Fia. 244. — Curve showing how the air is raised in temperature as it passes along the ducts and

the number of kilowatts absorbed.

the difference of temperature between the iron surface and the air, the velocity

of the air and the watts absorbed. If we take the slope of the curve giving the

watts absorbed, say at a point 16 cms. from the entrance to the duct, the slope

of this line gives us the kilowatts absorbed by the air per centimetre travel. At

the point of the 16 cms. the rate is 800 watts per centimetre. The temperature

difference between the air and the iron at this point is 9°, and the total area of

the ventilating ducts to which the air is exposed in traversing the centimetre length

of path is :

(300 X 2 X 21) + (72 X 2 X 21) = 15,600 sq. cms.

If we denote by A„ the watts per square centimetre of cooling surface per ° C.
difference of temperature between surface and air, we have

800

K = ^

= 00057.

' 9x15,600
This is at an air velocity of 3-95 metres per second.

250

DYNAMO-ELECTRIC MACHINERY

In order to see the effect on the distribution of temperature throughout the
machine with a greater draught, in test (b) the air supply was increased to 8800
cub. ft. per minute, the iron loss and excitation losses being as before. The tem-
perature of the air in various ducts is given in Fig, 245, and the temperature of

//? of i/e/f^/Az^//rg' afuc^.

Fig. 245.— Rise in temperature of air with air supply doubled and losses as before. TMt (B).

the various parts of the surface of the ducts in Fig. 246. Plotting the average
values of the air temperatures at different depths in the ducts, we get the curve
marked " Temperature rise of air " (Fig. 247), and plotting the average values of
the surface temperatures of the iron, we get the curve marked " Temp, rise of iron "
(Fig. 247). On this Fig. is also plotted the velocity of the air as it passes along the
ducts and the kilowatts absorbed by the air. Taking the tangent of the watts

d / Jj $~3 67 h a io }t ieaM^^/oirtsaAz/iiBS

Af9 of venC/kLb/ng^ ofucb.
Fig. 246.— Rise in temperature of iron with air supply doubled and losses as before. Test (&).

absorbed at the point 16 cms. from the internal cylindrical face of the stator, we
find that the air is picking up heat at the rate of 1250 watts per centimetre length
of path. The difference of temperature between iron and air at this point is 7'' C,
and the total area of surface exposed for 1 cm. of path is 15,600 sq. cms. as before ;
we therefore have

the velocity of the air being 7 -9 metres per second. We see from these experiments,
therefore, that A,, (the watts per square centimetre of cooling surface per ° C.

THE PREDETERMINATION OF TEMPERATURE RISE 261

difference of temperature between surface and air) is almost exactly proportional
to the velocity of the air. A,,, in fact, is given by the equation :

Jh =000145t7,

where t; is the velocity of the air in the ventilation duct in metres per second (see
Kg. 248).

e^oo cubic feet of Air per min.
/33 amperes eoccibd^ion*

O £

<f 6 a iO /^ n Id fS 20 Z2L B/i- i6 IS

D/c6(ince from dfrgdp in centimetres.

RG. 247.

In test (C) the air supply was maintained at 8800 ft. per minute, but the iron
loss was increased to 56 k.w. and the excitation losses to 17*5 k.w. Under these
conditions the temperature distribution of the air and iron in the ducts is given
by Figs. 249 and 250 respectively.

Ctonductivity of iron puncMngs. If we have a packet of iron punchings in
which the loss per cubic centimetre is constant, and if all the heat generated is con-
ducted across the packet and given off synmietrically to the air in the ventilating
ducts which bound it on each side, the hottest part of the punchings will be in
the centre, and the temperature gradient at any point within the iron will be

252

DYNAMO-ELECTRIC MACHINERY

proportional to z, the distance of the point from the centre. Let w be the watts
lost per cubic centimetre, then wdx will be the loss in a little part of the iron

Fio. 248. — Relation between A», the watts per square centimetre per * C. (difference in temperature
between iron and air), and the velocity of air in the ventilating duct.

O 7 1 3 ?~J 67 2r"5 M // /£ /3 /^ JSAS // /O /9 iO &I ££ £i%

Af? of i/enMidZ/T^ c/ucC.
Fio. 249.— Rise in temperature of air with supply doubled and iron loss increased to 56 K.w. Test (o).

C 7 £ ^ ^ S 67 6 a A> // /SS5/fi5/6/7/dJSJOJUat

A? cf ventfhtf/ig' duot.

Fig. 250. — ^Rise in temperature of iron with air supply doubled and iron loss increased to 56 K.w.

Test (0).

laminations 1 sq. cm. in area and dx cms. thick. The total heat generated in a
block 1 cm. high and 1 cm. wide, and of length x will be wx. If K,, is the heat
conductivity in watts per square centimetre per ° C, difference of temperature

THE PREDETERMINATION OF TEMPERATURE RISE 253

per centimetre, the temperature gradient -^ multiplied by the heat conductivity

d6
is equal to vxc. As -y- is negative when x is positive, we have

w
6 = constant - jr^ x^.

The curve of temperature distribution within the iron is therefore a parabola
such as that plotted in Fig. 243.

In the experiments above described, measurements were made of the tempera-
ture in the centre of the packet and on the exterior. Knowing the loss per cubic
centimetre of iron, we can calculate Kj^ as follows :

Example 41. The total iron loss amounted to 43*5 k.w. Of this 4*5 K.w. was in the
teeth, and 39 K.w. behind the teeth. The total volume of iron behind the teeth was 710,000
•cub. cms.

39 000
Thus, ' ■! /yw^ = 0'055 watt per cubic centimetre = ir.

It is seen from Fig. 243 that the temperature on the medial line between P and ^ was almost
constant for points on both sides of 0, so that the amount of heat conducted from 0 towards P
and N would not be very great. It would not be negligible, because the conductivity of the
punchings in this direction is so much greater than the conductivity across the laminations.
Let us take the figures at the point O, where the heat conducted along the laminations is at a
minimum, and calculate the conductivity on the assumption that all the heat flows to the walls
of the ventilating ducts.

Now there are 8° difference of temperature between the centre and the surface of the packet,
so we have

38 = 46 - -2p^ (see Fig. 243).

8=^(2-2o)«.

Kh= 0*0174 watt per square centimetre per "C. per centimetre.

Kk= 0*0042 calorie per second per square centimetre per ** C. per centimetre.

The formula given * by Dr. Ott for the heat conductivity across laminations is

Ki K^ tt
where

6] = thickness of iron in centimetres.

6^ = thickness of insulation in centimetres.
K^ — conductivity of iron = 0*15.
K^ = conductivity of insulation (paper = 0*0003) (varnish = 00006).

a = conductivity of rough surface. This may be between 0*5 for smooth and
0*04 for very rough iron.

In our experiments 5^ = 0041, ^2 = 00033, the insulation being paper. The
formula gives ^t = 00035.

*Ludwig Ott, Mitt. u. Fornchungsarhtiten, Heft 35 and 36, p. 63. See also T. M. Barlow,
"Heat Conductivity' of Iron Stampings," Journal of the Institution of Electrical Engineers,
vol. 40, p. 601; R. P. Gifford, "Influence of Various Cooling Media upon the Rise of
Temperature of Soft Iron Punchings," ibid. vol. 44, p. 753 (1910

254 DYNAMO-ELECTRIC MACHINERY

In the experiments described the loss per cubic centimetre, 0*055 watt, was
rather high. This was because the machine was run at 30 per cent, above its
normal field excitation. A more usual figure for 50 cycles would be 0*045 watt
per cubic centimetre. If then we take the conductivity of the punchings at 0*0174
watt per square centimetre per ® C. per centimetre, then we have 6, the difference
in temperature between the surface and middle of the packet

0*045 2
"■ 2 X 0-0174* •

For a packet 4*5 cms. thick (a; = 2*25) the excess of temperature would be 6*5** C,
and the mean temperature of the iron above the surface only 4*5° C.

At 25 cycles the loss per cubic centimetre would be about 0*025 watt per cubic
centimetre. Here the packets might be about 6 cms. thick for the same tempera-
ture rise in the hottest part.

In any case, it is seen that, unless the packets are made much thicker than is
usual in practice, the temperature rise in the centre due to the poor heat con-
ductivity across the laminations is not of very great importance.

Coolmg of external surface of staters. The cooling of the iron of a stator is
considerably helped by the conduction of the heat into the cast-iron frame, from
which it passes by radiation and convection to the surrounding air. On slow-
speed machines on which the depth of punchings is usually small compared with
the depth of the frame, this cooling by conduction is of more importance than on
turbo-generators with very deep punchings. It is in general impossible to make
an accurate calculation of the amount of this conduction, and yet one must make
an allowance for it in machines with shallow iron. Perhaps the simplest rule,
and one which gives a result not very far from the truth, in machines of normal
construction, is to allow 0*15 watt per sq. cm. for the whole of the external surface
of the punchings. This allows a temperature rise of 40** C. above the air. By
external surface we mean the external circumference of the punchings multiplied
by the gross length plus the area of the end plates flanking the iron at both ends.

Example 42. In the 750 K. v.a. generator illustrated in Fig. 3.32 we have a bore of frame of
184 cms. and a length of 31*8 cms. So the area of the bore is 18,000 sq. cms. The total area
of the flanks of the iron on both ends is 15,000 sq. cms., giving altogether 33,000 sq. cms.
Multiply by 0*15 watt per sq. cm., and we get 4950 watts conducted and radiated from the
outside area.

Collection of roles for predetermininir the cooling conditions. We may then
collect our rules for ensuring the cool running of a machine as follows :

1. Sufficient air must be provided to carry away the heat generated. A
supply of 100 cub. ft. of air per minute will in general be sufficient (see pages 206,
216 and 245).

2. Sufficient cooling surface must be provided to communicate the heat to the
air, and the short rules given below will in general tell us how much surface to
provide.

3. For ventilating ducts we may take the formula

hv=Krt\
where hv is the watts per sq. cm. of cooling surface per ° C, and Kt, is a coefficient

THE PREDETERMINATION OF TEMPERATURE RISE 265

between -OOOS and 'OOH (see p. 242). The difference in temperature between
surface and air and v is the mean velocity of the air in the duct in metres per
second. Where the machine is enclosed and provided with a definite amount
of air V is known. In other cases it should be roughly estimated from the
circumstances. A rule which works well enough in practice is to take v in the
ducts at one-tenth the peripheral velocity of the machine (see page 325).

4. For the cooling of the surface of rotors and the internal cylindrical face of
stators we may take the formula

watt* per sq. cm. _ 1 +0-1 v /, x

^Hse" "~333 ' ^ ^

where t; is the peripheral velocity of the machine in metres per second.

5. To find the number of watts conducted to and dissipated by the external
frame for a temperature rise of 40° C, multiply the " external surface '* (see page 325)
by 0 15 watt per sq. cm.

6. To find the difference of temperature between an armature coil and the
surrounding iron, one can adopt the method given on page 224, using the constants
for the heat conductivity of the insulating material given in Table XIII., and
allowing for air-spaces whose i^sistance is given roughly by Fig. 226.

7. To find the temperature rise of the surface of wire-wound coils upon which
the air is blowing with a velocity of v metres per second, we may take the formula

Ad=0-0011(l+0-54t;2) (2)

8- To find the difference between the inside temperature of a wire-wound coil
and the external temperature, we may follow the method given on page 236.

9, To find the difference between the temperature of the centre and the cooler
parts of a hot-bed of conductors cooled mainly by the conduction of the heat along
the conductors, we must adopt the method given on page 227.

10. To find the watts dissipated by the surfaces of a revolving field-coil, we
may adopt the rules laid down on pages 232 to 235.

Examples of the application of these rules in actual cases will be found on
pages 323, 349, 389, 454, 492 and 545.

The articles* referred to below bear upon the subject under consideration,
and will be of service to the reader.

* " Heating of Electrical Machines," Goldschmidt, Eltktrot. ZeU, 29, pp. 886, 912, 936, 1908 ;
" Heating of Armatures of Electric Machines," G. Schmalz, Elektrolechn. Zeitichr., 29, p. 188,
1908 ; " Heating of Ventilated and Enclosed Motors," Hartnell, Inst. E.E. Joum,, 41, p. 490,
1908 ; " The Heating of Induction Motors," A. M. Gray, Amer. I. E.E. Proc. 28, p. 606, 1909 ;
" Heating of Armatures," G. Ossanna, Elektrot. u. Maachinenbau, 27, p. 489, 1909 ; " The
Heating of Dynamos," £. Boulardet, Rev. Electrique, 16, pp. 608 and 662, 1911 ; " The Heating
of Electric Machines," C. Caminati, Lumiere Electr., 16, p. 147, 1911 ; '* Heating of Electrical
Machinery," E. Hinlein, Zeitschr. Vereines DeuUch. Ing., 66, p. 730, 1911 ; "Effect of Room
Temperature on Temperature-rise of Motors and Generators," Day and Beekman, Amer. Inst. E.E,
Proc., 32, p. 415, 1913 ; " Effect of Air Temperature, Pressure and Humidity on the Temperature-
rise of Electric Apparatus," Skinner. Chubb and Thomas, Amer. InM. E.E. Proc., 32, p. 563, 1913 ;
"Internal Heating of Stator Coils," Williamson, Amer. Inst. E.E. Proc., 32, p. 437, 1913 ; "In-
fluence of the Cooling Medium on Temx)erature-ri*ie of Stationary Induction Apparatus," Frank
and Dwyer, Amer. Inst. E.E. Proc., 32, p. 337, 1913.

256 DYNAMO-ELECTRIC MACHINERY

PERMISSIBLE TEMPERATURES.

The temperature at which electrical machinery may run for long periods of
time without sufiering injury depends upon the character of the insulating materials
used in its construction. The Sub-Committee on Rating of the American Institution
of Electrical Engineers have divided the insulating materials into the following
classes :

Class.

A L Fibrous materials which have not been specially treated for the purpose
of increasing their mechanical strength or durability under high
temperatures, such as cotton, paper and fibre. As a rule, such
materials become brittle or lose their fibrous strength when subjected
for a long time to moderately high temperatures.

A 2. Fibrous materials which have been subjected to a filling treatment with
oil, gum, or similar substance which increases their mechanical
resistance to disintegration. Impregnated cotton or paper fall into
this class when the filling compound has not been so applied as to
exclude air.

A 3. Fibrous materials which have been impregnated and " solid-filled " so
as to exclude air.

B 1. Those insulations which consist mainly of mica or asbestos, cemented
or impregnated with synthetic resins or other like material, whose
presence fixes the permissible temperature, but in which the air
is not excluded from the windings where they are employed.

B 2. Preparation of mica, micanite or asbestos applied to windings which
are " solid-filled " with impregnating compounds so as to exclude air.

C. Fireproof materials, such as pure mica, porcelain, etc., for which no tem-
perature limits are specified.

It is suggested that the highest temperatures to be permitted in a machine, in
any part insulated with these materials, shall not exceed respectively the following :

Class A 1 - - - - 90°C.

A2 - - - - 100°.

A3 - - - - 105°.

Bl - - - - 125°.

B2 - - - - 130°.

C - - - - No temperature limit specified.

It would seem logical that the rating of a machine should be so fixed that when
running continuously at its normal rating, or for a short period on overload, these
temperatures should never be exceeded.

" Observable " temperature. Where temperature is measured by thermometer,
or by the means ordinarily employed in temperature tests, it is very rarely possible
to find the temperature of the hottest part. The highest measurable temperature
will usually be smaller than the highest temperature attained in any part. The

THE PREDETERMINATION OF TEMPERATURE RISE 257

highest temperature measured may be termed the " observable " temperature.
The " observable " temperature will be less than the highest temperature attained
in any part of the insulation by the amount of internal temperature drop. In
practice, this internal temperature drop cannot be measured, but it can be arrived
at approximately by the application of data which have been obtained by scientific
investigation. One may ascertain approximately the highest temperature reached,
by adding to the observable temperature a suitable number of degrees for the
internal drop. The factors which control the amount of the internal drop are those
which have already been considered in this chapter, p. 236. The American
Institution of Electrical Engineers Sub-Committee suggest * the following approxi-
mate figures for the internal drop as a fimction of the voltage of the machine :

Up to and including 4000 volts - 10** C.

Above 4000 volts, and not exceeding 14,000 volts - - 20"* C.

Thus they arrive at the following observable temperatures of winding for the
rated pressures stated at the heads of the vertical columns :

In windings of rotating

apparatus, preeeures up to and

incluafng 4000 volts.

In windings of rotating

apparatus, all pressures

between 4000 and 14,000 volts.

(90-10 = ) 80°

(90-20 = ) 70°

(100-10 = ) 90°

(100-20 = ) 80°

(105-10 = ) 96°

(105-20 = ) 85°

(125-10 = ) 115°

(125-20 = ) 105°

(130-10 = ) 120°

1 (130-20 = ) 110°

Permissible temperature rise. Logically, the permissible temperature rise
can only be fixed if we know the temperature of the surrounding air in which the
machine is intended to work. Thus, for a machine intended to run in a sub-station
in a temperate climate, where the temperature of the surrounding air does not
rise above 25° C, an actual temperature rise of 65°, or an observable temperature
rise of 55°, would be permissible. On the other hand, where a machine is intended
for a tropical chmate, to work in a surrounding atmosphere which may at times
reach 45° C, the observable temperature rise ought not to be more than 35° under
the heaviest conditions of load.

In cases where the coils of a machine are very bulky, or, for any other reason,
have a considerable temperature drop between the interior and the exterior (see
page 236), the observable temperature rise should be even smaller.

* At the time of going to press, the result of the deliberations of the sub-oommittees of the
International Electrotechnical Commission, the British Engineering Standards Committee, and
the British Electrical and Allied Manufacturers' Association, had not been published. The
consensus of opinion is, however, in general agreement with the suggestions of the Sub-
committee of the American Institution of Electrical Engineers. The temperature of shunt
windings should be ascertained by increase of resistance ; the temperature of armature windings
may be obtained either by increase of resistance or by thermometer, and in the case of
measurements by thermometer the temperature recorded should be 5 per cent, below that
permissible when measured by rise of resistance. In the case of insulation of class B 2, the
temperature as ascertained by increase of resistance should not be more than 115° C.

W.M. R

PART II

THE SPECIFICATION

AND

THE DESIGN TO MEET THE SPECIFICATION

PART IL

CHAPTER XL

THE SPECIFICATION AND THE DESIGN .TO MEET THE SPECIFICATION.

Having given in Part I. a general statement of the properties of the materials
used in construction, and the rules which lead us to certain shapes and dimensions
in the design of Dynamo-Electric Machinery, we will now consider the form of
the specification which prescribes the performance of machines intended to be run
under certain conditions. We will then proceed to apply the rules set out in
Part I. to work out the details of machines designed to meet given specifications.

For each type of machine, whether it be A.c. or c.c. generator, induction
motor or rotary converter, there will be many different circumstances arising
in connection with the purpose for which it is used, which will lead to the specifica-
tion of definite qualities in the machine.

Given the duty that has to be performed, certain qualities should be called for
in the specification, and it will be part of our business in this section of the
book to show how the specification should be worded so as to describe what is
wanted, without interfering with the province of the designer and the manufacturer.

Then, given a certain specification calling for definite qualities in the machine,
another part of our duty will be to show how the manufacturer might design the
machine so as to give it those qualities in the most economical and satisfactory way.

It will be convenient for this purpose to take each class of machine in order,
and jconsider two or three machines in each class intended for work calling for
widely differing characteristics, and after giving the purchaser's specification
in each case, to work out a design fully to meet that specification. But, first, we
will make some general remarks upon the purchaser's specification.

PERFORMANCE SPECIFICATIONS IN GENERAL.

Main object of specification. A purchaser's specification of a dynamo-electric
machine should aim mainly at stating the duty that it is intended to perform,
the conditions imder which it will operate, and the tests that will be applied in
order to ascertain whether the performance is satisfactory. It should leave to the
manufacturer considerable licence to adopt such methods of construction as he inay
prefer, in order to obtain a machine which shall fulfil the prescribed conditions.

262 DYNAMO-ELECTRIC MACHINERY

For instance, it is much more important to state that a machine is intended to
operate in an engine-room having a temperature of 110° F. in a damp climate
than it is to specify " that the coils shall consist of copper wire having a conduc-
tivity not less than 98 per cent, of Matthiesson's standard," " that the current
density in the conductors shall not be more than 1500 amperes per square inch in
the armature coils and 2000 amperes per square inch in the field coils," and " that
the armature shall be built up of thin laminations of Swedish iron."

The manufacturer, for his own protection, will use copper of high conductivity
(generally of 100 per cent, of Matthiesson's standard). Copper of high conductivity
is very ductile and less liable to break when bent around corners than copper of
lower conductivity. The clause as to conductivity comes down to us from the
early telegraphic cable days, when it was necessary to instruct the manufacturer
in his art. Then, again, the current-density at which it is advisable to work the
copper in any part of a machine depends largely upon the cooling conditions. It
will often be found that in shunt coils the manufacturer cannot work the copper
higher than 800 amperes per square inch if he is to meet the temperature guarantees,
while in some parts he may, with impunity, employ 3600 amperes per square inch,
and yet give a thoroughly cool and satisfactory machine.

Of course, where the purchaser has a preference for some particular type
of construction, or for the use of a particular grade of material, it is important
that he should state his preference, always giving the manufacturer the opportunity
of substituting some other construction or material, which he can demonstrate
to be better adapted to the machines as manufactured by him.

The purchaser is interested in the qualities of the materials used in so far as
those qualities affect the permanent character of the work. Thus he may usefully
specify the tensile strength of materials employed, or object to certain methods
of insulation which experience has shown to be treacherous.

Arrangement of clauses. It is important that the specification should have
its clauses arranged in such a manner that matters of the same character are dealt
with together and in natural sequence. It often happens that in the perusal of a
specification by the staff of a manufacturing firm, different parts of the specification
are dealt with by different individuals, and it therefore contributes not only to good
feeling on the part of these individuals, but also to the efl&ciency of the specifica-
tion in stating the matter in hand, if each man who reads it, finds the parts with
which he is concerned without having to read through all the clauses.

A good way to begin is to state the kind of machine required, whether generator,
motor or rotary converter, and then give in tabular form the rating, so that any-
one can see at a glance the size and character of the machine required. For
instance, in specifying an A.c. generator, one might put the data of its rating
in tabular form as follows :

ENGINE-DRIVEN ALTERNATING-CURRENT GENERATOR.

Normal output 1250 k.v.a., 1000 k.w.

Power factor of load - . . . 0'8.
No. of phases 3.

SPECIFICATION AND DESIGN TO MEET SPECIFICATION 263

Normal volts 6300.

Voltage variation 6000 to 6600.

Amperes per phase - - - - 115.

Speed - - 250 revs, per min.

Frequency 50 cycles per second.

Kegulation 8 per cent, rise with non-inductive

load thrown off.

Over-load 25 per cent, for 2 hours, and 50

per cent, for 15 minutes.
Exciting voltage 125.

Temperature rise after six hours' full load 40° C. by thermometer.

50° C. by resistance.

Temperature rise after two hours' 25 per 55° C. by thermometer.

cent, over load 65° C. by resistance.

Puncture test 13,000 volts (alternating) applied

for 1 minute between armature

coils and frame.

1000 volts (alternating) applied for

1 minute between field coils

and frame.

This list of particulars is not intended to give full information ; it is merely
intended to give at a glance the general rating of the machine required.

Or, if the machine in question were a rotary converter, the principal data
might be set out as follows :

ROTARY CONVERTER.

Normal output

Number of phases - . . -

Frequency

Continuous-current voltage -

„ ,, amperes - - - -

Running A.c. to c.c - . - -

„ c.c. to A.c. . - - .

Compounding

Adjustment of voltage on rheostat
H.T. power factor at 550 volts full load
Over-load

Temperature rise after six hours' full load

Temperature rise after three hours' 25

per cent, over load
Puncture test

500 K.w.

50 cycles per second.

550.

910.

Yes.

Y^es.

530 to 550.

500 to 550.

0*97 leading.

25 per cent, for 3 hours.

50 per cent, for 10 minutes.

40° C. by thermometer.

50° C. by resistance.

55° C. by thermometer.

65° C. by resistance.

1500 volts (alternating) for 1

minute between windings and

frame.

264 DYNAMO-ELECTRIC MACHINERY

It is then convenient to make a general statement of the purposes for which the
machine is required, such as the nature of the load to be supplied, the machines
(if any) with which it is necessary to run in parallel, the location of the power
house or other running position, and any circumstances which may make the
performance difficult.

Then may follow clauses giving fuller particulars of the electrical rating, such
as will be found in the model specifications given in this book.

Finally, care must be taken to state exactly how much the specification is
intended to cover, in the matter of erection and setting to work, in the matter of
foundations and the provision of connecting cables and auxiliary appliances.

CHAPTER XII.

ALTERNATING-CURRENT GENERATORS.
HIGH-SPEED-ENGINE TYPE.

ALTERNATiNG-cintRBNT generators differ somewhat in their construction, according
to the service for which they are intended. In the first place, the prime mover
employed may be any of the following : a slow-speed engine (perhaps a gas engine),
a high-speed steam engine or motor, a water turbine, or a steam turbine, and it
will be necessary to adapt the design of the machine to speeds which are suitable
for these various prime movers. Secondly, the kind of load will vary in different
cases. We may have an intermittent motor load of low-power factor, with or
without lighting in parallel, or we may have a steady lighting load, or we may be
called upon to deliver current at a widely varying voltage, as to an electric furnace.
The size of the other units running in parallel and the size of the general system of
distribution will also influence us in prescribing the characteristics which the
generator under consideration must have.

We will therefore consider here four well defined types of alternate-current
generators :

(1) A 750 K.V.A., three-phase, 60 periods, 2100 volts generator, designed to be
driven by a steam engine running at 375 B.F.M., and to supply a load consisting of
induction motors varying in size from 10 to 100 h.p. with 50 K.w. of lighting load
in parallel.

(2) A 2180 K.V.A., three-phase, 50 periods, 6300 volts generator, intended to be
driven by a gas engine at 125 b.p.m., and running in parallel with similar machines
supplying a lighting and traction load.

(3) A 2500 K.v.A. generator driven by a water turbine at 600 B.P.M., and
generating three-phase current at 6900 volts, which is to be transmitted over a line
to various sub-stations for the supply of municipal lighting and power.

(4) A 15,000 K.v.A. machine driven by a steam turbine at 1500 R.P.M., and gene-
rating three-phase current at 11,000 volts for general municipal supply.

In each of these cases we will consider the characteristics which the generator
should have, and then give a suitable specification. We will then work out a
design of a generator to meet the guarantees asked for in each specification. In
connection with each design, it will also be possible to consider what variations
might be made to suit possible variations in the conditions.

266 DYNAMO-ELECTRIC MACHINERY

There are one or two matters affecting all A.c. generators that should be dis-
cussed before we pass on to the individual specifications.

Regulation. The inherent regulating quality of the machine to be asked for
will depend upon the character of the load and upon the number and output of the
generators in the power house. Where the output of the station is not large, and
the load is unsteady, a machine of fairly good regulation will be specified. But
where the changes in load are small compared with the total output, a cheaper
machine of poorer regulation may be specified. For a mixed power and lighting
load, in which the lighting is of first importance, it is usual to install an automatic
regulator. Even a machine of 6 per cent, regulation is hardly steady enough
when motors are started. In cases where the lighting is of secondary importance,
it is quite common practice to install a machine of such inherent regulating quali*
ties as to give 8 per cent, rise in voltage between full-load unity power factor and no
load, and, say, 22 per cent, rise in voltage between full-load 0*8 power factor and no
load. Such a machine, if installed without an automatic regulator, might com-
monly show on the voltmeter a drop of 10 per cent, to 16 per cent, when large
motors are started up. This voltage drop would not be of great importance if we
are not concerned with the lighting. If, however, it is important to keep the
incandescent lamps steady, a regulator of one of the well-known types will be
installed. Where a regulator is used, a generator of somewhat poorer regulation
than that mentioned is often installed ; but in view of the fact that no regulators
are instantaneous in action, or are immune from getting out of order, it is better
practice to install a machine of fairly good regulation — say with not more than
8 per cent, rise in voltage between full load and no load, and 12 per cent, drop in
voltage between no load and full load on unity power factor. Where the generator
is to work alone, an even better regulation will be preferred by some users, if it
can be obtained at a reasonable price, and there is no doubt that the operation of
the plant is somewhat more satisfactory. A generator of good regulating qualities
can, without much additional cost, be given a very much greater over-load capacity
than a machine of poor regulation. Many cases have arisen in which good regulating
generators have had their armatures rewound, and the capacity increased two-
fold. Where the load on a station has increased beyond expectation, the fact that
ample machines were originally installed had resulted in a great saving of money.
Where three or four generators of fair size are working in parallel, and the large
induction motors are all started up on resistances so that they do not make heavy
demands for wattless current, it will be found that generators giving not more than
25 per cent, rise in voltage when full inductive load is thrown off are in general
satisfactory. It is not very important that all the generators running in parallel
shall have the same regulating qualities. The only effect of running a good regula-
ting machine in parallel with a poor regulating machine is that the good machine
tends to take more than its share of the wattless current as the load increases.

Temperattire rise. The main object to be kept in view in specifying tempera-
ture rise is to ensure that in the ordinary course of operation no part of the machine
shall attain a temperature which will permanently injure it. It is, of course,
impossible to actually ascertain the temperature of internal parts, and one can
only use judgment based on past experience. Generally, it may be stated that

ALTERNATING-CURRENT GENERATORS 267

where the insulation is thick, as in high- voltage machines, there will be a tendency
for the temperatures attained by internal parts to be much higher than those
measured by thermometer. In machines of 11,000 volts, one may assume that
with ordinary methods of construction the temperature on the inside of an arma-
ture coil will be about 20° or even 25° higher than the tei!nperature of the sur-
rounding iron ; whereas in most low-voltage machines there will commonly be
not more than 10° difference in temperature between the inside of the insulation
and the surrounding iron. The internal layers of wire-wound coils (see page 236)
are often very much hotter than one might commonly suppose from a measure-
ment of the average temperature by the increase of resistance method. Where
the kind of machine {e.g. a low-frequency alternator of high speed) is auch as to
employ bulky field-coils, special care will be employed in the specification to prevent
excessive temperatures in the internal layers. There is some difference of opinion
as to how high a temperature such insulating materials as cotton and paper will
satisfactorily withstand for long periods of time. Although the cotton covering
of wire may be subjected to a temperature of 125° C. for long periods without being
destroyed, there is no doubt that any temperature over 100° C. will dehydrate
the cellulose in time, and render it extremely brittle. In parts where the con-
struction is such that movement of the conductors relatively to each other may
occur as in revolving armatures, the temperature should not be allowed to exceed
100° C, but in stationary coils not subjected to vibration the internal temperature
often exceeds 120° C. without any apparent harm. If we take 90° C. as a safe
temperature for cotton-covered wires of a revolving field coil at normal loads, we
can arrive at the allowable temperature rise measured by thermometer as follows :
Deduct from 90° C. the nimiber of degrees — 10 to 25 — by which the actual tem-
perature may exceed the measured temperature, and then from the remainder
deduct the temperature of the air which we expect to find in the power house.
For instance, the part of a 600 K.w. a.c. 440 volt revolving-field engine-type generator
which is most likely to deteriorate from excessive temperature is the cotton insula-
tion on the field coils. The inside of a field coil of a machine of the above rating
might be 20° C. hotter than the temperature measured by thermometer. This
is assuming square wire coils of about V winding depth (see page 236). Deduct-
ing 20° from 90°, we get 70°, and assuming the temperature of the air in the power
house to be 25° C, we may safely allow 45° temperature rise in the field coils
measured by thermometer. If the temperature were measured by the resistance
method, we might safely allow 55° C. rise.* It will be seen that where a machine
is intended to operate in a hot climate, and where the temperature of the power
house might for long periods be at 45° C, it would be well to specify 30° C. rise
for a low-voltage machine. For high-voltage machines, it might be justifiable to
call for a lower temperature rise. It must, however, be remembered that the
figure of 90° C. is rather on the safe side, and one might base one's figures on 100° C.
as the permissible temperature of the very hottest part. It depends on the rela-
tive importance of the reliability of the machine and its first cost (see page 256) .

* In revolving-field coils the ends of the coils exposed to the draught are rather cooler than
the parts between the poles, so that the average temperature of all the copper is not much
higher than the hottest spot that can be found by the thermometer.

268 DYNAMO-ELECTRIC MACHINERY

When a machine is run on 25 per cent, over-load, the losses in the armature
copper are increased about 60 per cent., but one does not ordinarily find that
the temperature of the armature copper is increased by 60 per cent., because the
losses due to air friction and iron loss do not increase appreciably with the load.

If a generator is of fairly good regulation, so that the field current is only increased
30 per cent, between no load and full load, the increase of field current will be
about 40 to 45 per cent, between no load and 25 per cent, over-load (see page 292).
That is to say, the field current may be increased 11^ per cent, above the full-load
value. We may expect a temperature rise 32 per cent, higher when at 25 per
cent: over-load, if we have an increase of the field current of llj per cent, and
an increase of the resistance 6 per cent., because 11 15x1 -115x1 06 = 1*32. So
we see that, if the field coil is the part which we expect to be hottest on over-load,
60° C. (as measured by thermometer) would be a reasonable temperature rise to
allow for this over-load. The temperature rise in the hottest part of the coil might
reach 87° C. above the atmosphere, if the over-load were maintained continuously.

Efficiency. In cases where an engine and generator are bought from one
contractor, the purchaser is generally not concerned with the efficiency of the
generator itself, so long as the steam consumption of the set per K.w. hour is
guaranteed. In many cases, however, the generator is sold by a contractor who
has no responsibility for the efficiency of the engine ; in this case the efficiency
guarantees are important.

It should be clearly stated how the efficiency is to be arrived at. Most manu-
facturers specify that the efficiency shall be calculated from the various losses
(copper losses, iron losses, etc.) measured separately. In this case, it should be
clearly stated what friction and windage are to be included among the losses.
Some makers will exclude all bearing losses when the bearings are supplied by the
engine builders. Others will include the losses in the outboard bearing. Where
there is a flywheel, the windage loss due to it will in general be included in the
losses of the maker who supplies the flywheel. In general, it is not good practice
for the purchaser to specify any particular efficiency. It is better to ask the con-
tractor to state his efficiency, calculated in a certain way. The losses to be included
should be clearly stated. The k.v.a. output of the machine, the power factor
of the load and the voltage at the terminals at which the machine is supposed
to be run when the efficiency is taken should also be stated. In allowing for increase
in the resistance of the conductors, it is sometimes assumed that they will reach
the temperature rise specified, though in cases where the copper is found on full
load to be very much below the specified temperature, the contractor is entitled
to take his " hot " resistances at the temperatures actually reached. This is of
special importance in the case of slow-speed engine-type generators having a very
great number of poles, because in these cases the temperature of the field is
generally fairly low, on account of the very large cooling surface (see p. 347).

Excitation. If there is always available in the power house a continuous-current
supply of a voltage not higher than 240, it is quite good policy to excite from this
supply, and an exciter may be added as a spare. The advantage of exciting the
field-magnet from a supply of constant voltage is that the exciting current is then
independent of the speed of the engine, and the regulation of the set is therefore

ALTERNATING-CURRENT GENERATORS 269

better. It is not, however, good practice to excite an alternating-current generator
from a 500 volt C.C. circuit (unless the output is very large), because the economical
size of wire to be employed is rather small for use on revolving field coils.

The most common method is to provide an exciter directly connected to the
end of the shaft of the main generator. This exciter will cost a little more than
a belted one running at a higher speed, but is generally considered more satisfac-
tory. Where an exciter is employed, the voltage chosen will generally be 125
volts. In some cases where an automatic regulator is installed, an exciter is a
necessary part of the equipment.

Bheostats. It is quite common practice to have no rheostat between the
exciter armature and the field-magnet, and to rely entirely upon a rheostat in
the field circuit of the exciter to obtain the necessary change in excitation. This
arrangement renders the regulation of the set much poorer than where a main
rheostat is installed, because the exciter voltage will change more with speed when
it is working with its field not fully excited than where it is working at full voltage.
The use of a main rheostat, of course, leads to some extra loss, which in the case
of a 600 K.w. generator at 375 r.p.m. would amount to be about 2 K.w. at half
load, if the exciter were always maintained at its full voltage.

High-voltage test. The purchaser's specification will state the testing voltage,
which will be applied between the armature winding and frame, and between the
field winding and frame. It will also specify the interval of time during which
the testing voltage is to be applied.

These matters may be in accordance with the rules laid down in Chapter VIII.
page 188. In case the working voltage in the armature is 2100, a suitable testing
voltage would be 5000 applied for one minute. If the field is excited at 125 volts,
a suitable testing pressure would be 1000 volts applied for one minute.

After these general remarks, we will proceed to make out a specification for
a 750 K.V.A. three-phase generator for 50 periods, 2100 volts, 375 R.P.M.

A machine of this character would, in all probability, be built on one of the
standard frames of the manufacturer; and if we wish to purchase a cheap
machine, it is desirable to avoid anything in the specification which will prevent
a manufacturer from quoting on his standard plant. The specification will
therefore, in this case, be as short as possible, and will not contain anything more
than is necessary to secure a generator which will satisfactorily perform the work
intended for it.

SPECIFICATION No. I.

760 K.V.A. THREE PHASE ENGINE DRIVEN GENERATOR.

1 . The work covered by this specification is to be carried Genej»i

•iii^ i/^T' iTfc 1 Conditions.

out m accordance with the General Conditions and Kegula-
tions issued by the Institution of Electrical Enjgineers, in so
far as they are not inconsistent with anything contained
herein.

270

DYNAMO-ELECTRIC MACHINERY

Extent of
Work.

2. The work includes the supply, deUvery, erection,
testing and setting to work on the site shown in the accom-
panying drawing No. of an alternating current generator
which shall have the characteristics set out below :

Characterifltlcfl
of Generator.

Normal output
Power factor of load
Number of phases
Normal voltage
Voltage variation
Amperes per phase
Speed
Frequency
Regulation

Over load

Exciting voltage
Temperature rise after
6 hours full load

Temperature rise after
2 hours over load

750 K.V.A., or 600 K.w.

0-8.

2050

2000 to 2100.

206.

375 revs, per minute.

50 cycles per second.

8 per cent, rise with non-induc-
tive load thrown off, the speed
and excitation being constant.

22 per cent, rise with 0-8 power
factor load thrown off, the
speed and excitation being
constant.

255 amperes at 2050 volts power
factor between 0-9 and unity.

120.

45° C. by thermometer.
55° C. by resistance.

55° C. by thermometer.
65° C. by resistance.

Running 3. Thc generator is intended to run in parallel with two

Conditions. /» • •! i i • n i

generators of similar output and speed at present installed
in a power-house, supplying power to a colliery, the most
Nature of Load, distant parts of which are about three miles away. The load
will consist of coal-cutters, three-phase haulage motors and
the lighting of the mine. The largest motors at present in-
stalled are 100 h.p., and are of the sUp-ring type. It is
proposed to install a 400 h.p. winding motor of the sUp-ring
type. The generator shall be suitable in every way for taking
this class of load.

Type of
Generator.

Connection to
Engine.

4. The generator shall be of the revolving field type, and
the spider shall be mounted in such a manner that it can be
very rigidly fastened to the flywheel of the engine. The
method of attachment shall be indicated in the outline supplied
with the tender.

ALTERNATING-CURRENT GENERATORS 271

5. With the generator the contractor shall supply a bed- Bedpute and
plate adapted for bolting to the engine bedplate, and an
outboard bearing. The bearing shall have a self-aligning
seating and be provided with approved means of adjustment.

6. The foundations will be suppUed by the purchaser Foundations,
to templates furnished by the contractor. The contractor

shall supply all foundation bolts. Within four weeks after
the acceptance of his tender, the contractor is to provide a
drawing showing the details of the bedplate and foundation
bolts, and the position of the terminals of the generator.
Cables from the generator terminals to the switchboard will cawes.
be provided by the purchaser.

7. The contractor is warned that the engine-room is in severe

T. 'jj* iji ■!• Ti i"! Conditions.

a dirty situation, and the machinery supplied must be
suitable to run under the existing conditions.

8. There is a railway track into the power house, and an Access to
overhead hand-operated crane capable of lifting loads of

10 tons from a railway truck to the proposed foundations.
The contractor may have the use of this crane at his own risk, crane.
and he shall be responsible for any damage done. The
tender shall state the maximum weight to be Kfted during
erection or overhauUng.

9. The generator shall run well in parallel with the existing Paraiiei
generators, which run well in parallel with one another.

10. The electromotive force wave form of the generator e.m.f. wave,
shall at full load be a smooth even curve* free from ripples or
pronoimced higher harmonics.

11. The generator shall be excited from the existing Excitation,
exciting bus-bar at 120 volts, and no exciter need be provided.

The exciting bus-bar pressure may at some future time be
controlled by an automatic regulator to maintain the a.-c.
voltage of the station constant, but the inherent regulation
of the generator must be as specified above, apart from any
automatic control.

12. The tenderer shall state in the tender what provision ^^/^"^
is made for obtaining access to the field magnet and armature

for inspection and repair. He shall also state the method
proposed of replacing armature coils and field coils in case
of a breakdown.

•See page 380 for a more stringent clause.

272

Short-circuit.

Permanent
Construction.

Oil-throwing.

Efficiency.

Kheoetat.

DYNAMO-ELECTRIC MACHINERY

13. The generator must be able to withstand a short
circuit at its terminak when running at full voltage, but the
contractor shall not be called upon to cany out a short
circuit test.

14. The contractor must be able to show by calculation
that the mechanical strength of all parts is such that when
running at full speed on load there is a factor of safety of
four. No part of the generator shall be of a material which
will deteriorate with time, and become so weak, brittle or
otherwise defective as to make the factor of safety less than
four.

15. The oil- throwing devices on the shaft and bearing shall
be so eflficient that no oil or oil vapour is apparent outside
the bearing during ordinary running without attention.
After erection and final adjustment, a special test shall be
made to see that this condition is complied with.

16. The efl&ciency shall be calculated in the following way :
The iron loss at 2100 volts and the friction and windage shall
be measured at no load. The armature resistance shall be
measured at a known temperature and the PR loss calculated
at 60° C. The field and rheostat losses shall be taken as
together equal to the number of amperes of field current
at 0-8 power factor multiplied by 120, the voltage of excita-
tion. All the above losses, expressed in kilowatts, shall be
added to the kilowatt output, and the ratio of output to this
sum shall be taken as the calculated efficiency. The con-
tractor shall state in the schedule attached the efficiency of
his generator calculated in this way at full, three-quarter
and half load on a power factor of 0-8, and he shall guarantee
that there shall be nothing in the construction of the machine
that will make the actual efficiency when running on load
more than 1 per cent, less than the figures so stated.

17. A field rheostat and multi-contact switch is to be
provided in the field circuit of the generator, of sufficient
capacity to lower the voltage of the armature to 1950 volts
at no load when the machine is cold. Sufficient contacts
must be provided on the switch to make the voltage change
very gradual as the switch is moved over the whole range.
One step of the rheostat must not change the voltage by more
than 15 volts at any load and at any part of the range when
the machine is operating by itself.

ALTERNATING-CURRENT GENERATORS 273

18. The slip-rings for the exciting circuit must be ofgjp^^j^^
sound metal free from blowholes and mounted in a manner

that ensures exact concentric running. The brush-holders
must be of a soUd, simple construction, rigidly supported,
and so made that the brushes can be easily inspected while the
machine is running. There must be at least two brushes per
ring (preferably at opposite ends of a diameter on each ring),
and there must be no heating or sparking at the rings with the
maximum field current flowing and one of the brushes raised.
The brushes must be of carbon.

19. The following tests shall be carried out on the gene-
rator :

(a) Measurements shall be. made of the resistances of Teats
the armature and field windings.

(6) The generator shall be run at ftdl speed at no load Ma«neti«ition
with the field excited, and measurements shall be taken
showing the relation between field current and voltage
generated, the iron loss at various voltages, and the
friction and windage.

(c) The generator shall then be run with the armature short-circuit.
short circuited, and measurements taken to show the
relation between the field current and the armature
current.

(d) From tests (a), (6) and (c) the field current re- Fieid-heating
quired at ftdl load 0-8 power factor shall be approxi- ^'
mately calculated, and the generator shall be run at

this field current for six hours, and measurements taken
of the field resistance while hot.

(e) While the machine is still hot an alternating Puncture Tests,
pressure of 5000 volts virtual shall be appKed between

the armature winding and frame for one minute,
and an alternating pressure of 1000 volts between the
field winding and frame for one minute.

(/) After erection on site or at the engine-builders' f J^"^^^^
works, as shall mutually be agreed upon, the generator »«n.
shall be run at full load, 0-8 power factor, for six hours,
and for two hours on the stated over load, and measure-
ments shall be taken of the temperature of the armature
windings and iron, and field windings, by thermometer,
and of the field windings by resistance, to see that the
specified temperature rises above the surrounding air
are not exceeded. For the purpose of these tests the

W. M. s

274 DYNAMO-ELECTRIC MACHINERY

temperature of the engine-room shall be taken three
feet away from the generator in a line with the shaft.

ueguiation. (^) If the puTchaser is not satisfied with the calculated

regulation figures obtained from tests (a), (6) and (c),
a regulation test shall be made on site after erection
by throwing off full load at 0-8 power factor to see if the
generator has the inherent regxiation specified.

Endurance. (^) After crcction on site the generator shall be run

on its ordinary daily load for one week under the
direction of contractor's engineer to see that all matters
are in order. It need not be accepted by the purchaser
until it is complete in every particular.

Spares. 20. The tenderer shall quote separate prices for the

foUowing spare parts :

(1) A field coil.

(2) Armature coils of various sizes.

(3) Brush gear and brushes.

(4) Bearing bush.

THE DESIGN TO MEET THE SPECIFICATION.

We will now consider the matter from the manufacturer's point of view. We
will suppose that he has received an order for a 600 k.w. generator, which is to
comply with the above specification. How can he most economically build the
machine ?

Ohoice of firame. The particular length and diameter of frame that he will
choose will depend upon what machines of similar size he has built before, and the
patterns and dies that he has available. It may be much cheaper for him to choose
a D^l much larger than the theoretical minimum than to build an entirely new
machine which shall employ the smallest possible amount of material. All that
we can do here is to choose a diameter and length which will be veiy economical
in material, and yet sufficient to enable a machine of simple construction to be
built which will safely meet the guarantees.

Depends on output of fleld-magnet. It will be found that with revolving
field A.c. generators of fairly good regulation the limiting conditions as to size lie
in the field-magnet. We must have a certain cross-section of steel in the poles to
provide the magnetic flux without undue saturation, and we must have sufficient
copper space for the field •ampere-turns at fuU load. At the same time, we must
provide air spaces for cooling between the field coils. These considerations determine
the size of the field-magnet. If we are sure that this is big enough, there will be
no difficulty with the diameter and length of the armature, because it will be found
that a reasonably shallow slot (2 to 2^ inches deep) will carry all the copper we

ALTERNATING-CURRENT GENERATORS 275

want in the armature, and there will not be much difficulty in providing sufficient
cross-section in the teeth to carry the magnetic flux, except in very high-voltage
machines.

We will consider first the field-magnet. It has been found that one of the most
economical constructions for high-speed engine-type generators having from 6
to 20 (or even more) poles is one consisting of a cast-steel spider with poles of mild
steel machined out of the solid bar and bolted on. For very high speeds, as for
water turbine-driven generators, the poles may be dovetailed in.

Gast-steel poles are sometimes used, but they are not always free from blow
holes. If the sides of the poles are not machined, a considerable amount of space
must be allowed on the inside dimensions of the field coils to allow for rough-
nesses of the casting. This is a bad feature. If the poles are machined on the
sides they will cost as much as, or more than, poles cut out of solid mild steel.

Punched poles. Where a manufacturer has a pole die of the right size, or
where the number of poles is so great as to make the cost of a new die of little
importance, he can build up poles out of punched steel just about as cheaply as he
can machine poles out of the soUd, for the overhanging horns of the pole necessitate
the machining away of a large quantity of metal.

The punched poles have the advantage that they can be used with open slots
in the armature without causing so much loss in the pole face. Where open slots
are used, and particularly where the width of the slot is more than double the
width of the air-gap, laminated poles should be used, or at least laminated pole
shoes.

When a pole is built up of punchings, it is comparatively easy to provide it
with tunnels near the pole face for the reception of copper rods to form a damper,
or to make slots in the iron to cause any required amount of saturation, or other-
wise make the pole of a complicated shape that might be expensive to make out of
solid metal. On the other hand, when once a die is made, we are to a great extent
restricted in our design by the shape of the punching. We cannot, without expense,
for instance, narrow the pole to make room for more copper in a case where that
course might be advisable, or widen the pole in a circumferential direction to get
in more iron in cases where the saturation is rather high. We can, however, always
build up a punched pole to a greater axial length when we wish to increase the
cross-section of the pole body.

Relation of width of pole to pole pitch. In designing a pole die for a par-
ticular frame, it is of great importance to make the width of the pole body in
relation to the pole pitch such that we can get the most economical arrangement
of material for those generators that are most conmionly built on that frame. In
the first place, we need hardly say that there is a great advantage in using a pole
with overhanging lip, as in Fig. 234. This lip enables us to make the pole face
as wide as we like, while we have a free hand with the width of the pole body. The
lip also gives a good mechanical support for the coil. It will be found that there
is no advantage in making the pole arc greater than two-thirds of the pole pitch
except in high-speed machines, where the coils are difficult to support.

Relation of pole arc to pole pitch. The only object in widening the pole arc
is to reduce the ampere-turns on the gap, but it is better to make a short air-gap

276 DYNAMO-ELECTRIC MACHINERY

than to widen the pole arc unduly. The magnetic flux which comes from the pole
lips is out of phase with the flux in the centre ot the pole, and is therefore not very
effective in producing useful electromotive force, while all the flux that comes from
the pole requires so much cross-section of iron in the pole body.

The bringing of the lips on North and South poles near together increases the
magnetic leakage, and by taking away from the usefulness of the pole body may
diminish the output of the machine. Upon the whole, on 50 cycle engine-driven
generators a pole arc about 0*64 of the pole pitch will be found to be most
generally useful. If the corners of the pole are bevelled off as in Fig. 334, it
will be found that the fringing from the lips and sides of the pole brings up the
fleld-form constant K^ to 0*64 (see page 16). The electromotive force constant Ke
for a three-phase winding of full pitch, and not less than two slots per phase per
pole, would with this pole be 0*4. For a two-phase winding, the constant K^ would
be 0-315.

Width of the pole body. We now come to a most important consideration
affecting the output and performance of an a.o. generator. In the first place,
it Vill be observed that most manufacturers employ a pole body with parallel
sides. This is because it is so much easier to wind the coil for a pole with parallel
sides, and to slide it on, than to make a coil to fit a taper pole and hold it in posi-
tion. Nevertheless, the taper pole has some strong claims if we wish to get
the greatest possible output from a given amount of material.

The drawback to the pole with parallel sides is that if we make the pole body
immediately below the lips as narrow as we would like to make it, and thus get
plenty of room for copper, we will find that the width at the bottom is too small,
and the saturation of the iron, particularly at heavy loads, is too great. If, on the
other hand, we make the pole at the base as wide as we would like to make it, we
have more iron than we need at the top of the pole, and we are cramped in our
copper space. If we use parallel sides, we must make a compromise, keeping the
saturation within sufficiently safe limits, and yet getting as much room for copper
as we can.

, The best width for a parallel pole depends upon the number of poles. Where
a machine has many poles, the centre lines of adjacent poles are nearly parallel,
and this gives more room at the root than where, the number of poles being few,
the centre lines are inclined to one another. For machines with 12 poles of moderate
regulating qualities the pole body is usuaUy made about half the pole pitch. As the
number of poles is increased, the ratio increases from 0*5 to 0*6. For 8 poles the
ratio is often as low as 0*47 and for 6 poles as low as 0*45.

The above figures are based on the assumption that copper costs &^. per pound
and iron \d. per pound. If the cost of copper increased very much, it would pay
to reduce the copper space and to somewhat increase the size of the whole frame,
putting in more iron.

Relation between weights of copper and iron. To arrive at clearer ideas as
to how the output of a frame depends upon the amount of copper and iron in it,
we may consider one of the poles of the 16 pole case given in Fig. 234. For the
same peripheral speed and length of iron the output of the machine is proportional
to the number of poles, so that we may consider one pole by itself, and aim at

ALTERNATING-CURRENT GENERATORS 277

making the proportions between the iron and copper such as to get the greatest

output for a given cost.

We have

output of three-phase generator = volts x amps, x 1*73,

and &om equation (1), page 24, we have

volts =KeX Rpg X AgB X 10"® X conductors.
Therefore

output = (Z« X Rpg X AgB X 10"*) X (conductors x amperes) x 1-73.

Now Kg depends on the pole-arc, but not necessarily on the pole-body width,
so we can leave it out of account in the present discussion. The two important

factors are :

(1) AgB = " magnetic loading."

(2) Conductors x amperes = Zaia = " current loading."

Now, the magnetic loading AgB can be increased by increasing the width of the
pole body. Within the limits of practical design AgB will be roughly proportional
to the width of the pole body.

For machines having the same ratio between field ampere-turns and armature
ampere-turns, the factor ZaIa can be increased by increasing the copper space
of the field. The increase of ZaIa will not be quite proportional to the increase
in the copper space, because the cooling conditions are worse with increased depth
of copper, but for very small increases in the copper space we may, for the sake
of the present discussion, take ZaIa as proportional to the copper space.

Now consider Fig. 234. The half width of the pole is 3*125 and the mean depth
of the copper winding 1*25 inches. Suppose that we were to take O'l'' off the side
of the pole and utilize for copper the space gained. It is clear that the flux factor
would be reduced by a little over 3 %, while the copper factor would be increased
by 8 %. Even allowing for the cooling conditions being somewhat worse, it is
clear that the output of the frame can be increased by making the pole narrower
and using more copper. But at what cost ? The extra copper for field and
armature will cost, say, Sd, per lb., while the saving in iron is very little. A
more economical way of increasing the output would be to leave the copper weight
as it is, and to increase the width of the pole. This would mean using a larger
frame, but an increase of the iron by 5 % right through the machine would not
cost as much as an increase of the copper weight by 8 %. As a matter of fact,
the problem of arriving at the best proportion between copper and iron is so com-
plicated by various considerations, such as the cost of labour, the freights on foreign
shipment and the expediency of standardization, that it is impossible to get an
exact solution. It is, however, clear from the machines put on the market by the
most successful makers, that it does not pay to get the greatest theoretical output
from a frame of given size. There comes a time in the loading of the frame when
the money put into extra copper is better spent in increasing the siz^ of the frame.
The proportions given in Fig. 234 are not far from what is practically the most
economical arrangement.

Now, the winding shown on the pole will (when the axial length of the machine
is about 12") carry about 10,000 ampere-turns for 45** C. rise, the peripheral speed

278 DYNAMO-ELECTRIC MACHINERY

being 6000 feet per min. (see page 303). We may therefore take 10,000 ampere-
tuma as full-load ampere-turns on the pole. What shall we take for the no-load
ampere-turns ? That depends upon the inherent regulation asked for, and leads
us to some general remarks on regulation.

THE REGULATION OF A.C. GENERATORS.

One of the main considerations which determine the size and cost of an a.g.
generator is its quality of maintaining its voltage within narrower or wider limits,
commonly spoken of as its " inherent regulation.'* Before we can fix upon the
size of the frame upon which to build a generator of a given output and speed,
we must see within what limits it is required to maintain its voltage when the
load changes.

The most usual way of specifying regulation is to give the percentage rise of
voltage when the load is thrown off, the speed and excitation being kept constant.
We will speak of this as " regulation up." As the iron of the field-magnet usually
becomes saturated * at voltages a little above the noimal, much closer regulation
figures can be guaranteed when the regulation is specified in this way than when
the percentage drop in voltage, when the load is thrown on, is specified. The latter
characteristic we will speak of as " regulation down,"

From the user's point of view a machine with 8 % regulation down is a much
more satisfactory machine than one with 8 % regulation up, but the cost of the
first machine will be considerably higher, so that unless the load is of a very fluctuating
nature, and it is required to maintain the voltage fairly steady, independently of
the action of an automatic regulator, the purchaser will be content to take the
guarantee most commonly offered by manufacturers.

The methods of predetermining the regulation of an A.G. generator from the
design data and from no-load tests, have formed very fruitful subjects for dis-
cussion in our text-books, and in papers before learned societies. No method known
to the author is perfectly accurate when put to the practical test. AU methods
assume a sine-wave form for the armature current, and none take into account in
a perfectly satisfactory manner the bevelling of the pole face or the saturation of
the iron. Fortunately, in ordinary commercial design it is not necessary to pre-
determine the regulation of a generator very closely. So much variation occurs,
even between two machines built to the same drawings, that considerable margin
must be allowed if the regulation guarantee is to be met with certainty, and there-
fore a superfine method of calculation is out of place.

The method which we shall give here is one which has stood well the test in
practical manufacturing, and while probably as accurate as any other for generators
with cylindrical field-magnets,')' it is very easy to apply and to understand.

Consider a two-pole generator provided with a cylindrical field-magnet, such
as is generaUy found in high-speed turbo-generators. The field winding usuaUy
occupies some 75 or 80 per cent, of the circumference. This is indicated by the

* There are other matters beside the saturation which tend to make the regulation down
much wider than the' regulation up. These are considered later.

fThe case of generators with salient poles is considered on page 293.

ALTERNATING-CURRENT GENERATORS

279

dots and crosses on the inner circle in Fig. 300. A dot represents a current
coming towards the observer, and a cross a current going away. The regulating
qualities of the machine will depend to a certain extent upon the width of
the pole.

We will consider a three-phase generator, because this is the kind of generator
most commonly built, and its armature reaction behaves as a rotating vector
of almost constant value.

^a^aj^i^n

® ® ®

Fio. 300. — Showing the effect of aniuiture reaction on the strength of a cylindrical field-magnet.

Take the instant at which the current in phase A is at its maximum and going
away from the observer in the six slots at the top of the armature in Fig. 300. The
current in phases B and C will then be at one-half their maximum value, so the
ampere-turns of the armature tending to drive flux along a horizontal diameter
will be

1-41 X /a X

2p^3"

. After one-twelfth of a period has elapsed, the current in phase A will have
sunk to 0-866 of its maximum value, and the current in phase 0 will have risen to

280 DYNAMO-ELECTRIC MACHINERY

0-866 of its maximum value, while phase B will be at zero. At this instant the
ampere-turns on the armature will be

0-866 X 1-41 X /a X 2" X g =

Observe that we are only considering the magnetomotive force along the
centre-line of the magnetic path. The field form of these cylindrical magnets is
so nearly sinusoidal and the effect of the armature reaction is so close to what
would be produced by a sinusoidal distribution * of current that we find it sufficient
to consider only the crest values of the magnetomotive force (see page 396).

The mean value of the ampere-turns of a three-phase armature is therefore

0437/aZa,

where la is the virtual current per conductor and Za the total number of con-
ductors on the armature.

Let us represent these ampere-turns by a vector Iga drawn from 0 and point-
ing to the centre of the phase band A, The vector points in the direction in
which the current flows along the connectors from the bottom to the top
of the armature. The flux produced by these ampere-turns if acting alone
would be along a horizontal diameter, but it will be found more convenient
to draw the ampere-turn vector pointing to the centre of the phase band
A than to draw it along the line of the flux. In the same way, the ampere-
turns of the field-magnet can be represented by a vector I^ drawn parallel
to the direction of the current in the end connectors of the rotor. If we add
together the vectors I^ and /,«, we get the resultant vector Im which gives the
actual magnetomotive force on the magnetic circuit of .the generator. This will
create the actual working flux along a line at right angles to /»-. The centre of this
flux will be on the line of the vector Eg, and if we know the magnetization charac-
teristic, we can draw the vector Eg to the volt scale to represent the phase and
amount of the generated E.M.F. Here we have drawn the vector representing the
E.M.F., so that it points to the phase band in which the E.M.F. is at its maximum.
It will be seen that the vector summation of I^. and Iga has taken into account
both the demagnetization and the cross-magnetization effect of the armature. If
we know the voltage drop in the armature due to its self-induction we can set off the
reactance voltage by the vector laXa and the drop in the armature resistance by
the vector lafat ^^d thus we arrive at the terminal E.M.F. Et. This is, of course,
made up of the ohmic drop laR in phase with the armature current and the inductive
drop, in the outside circuit, represented by the horizontal vector laX.

Now, there are two ways in which this diagram may be arrived at. (1) By
calculation from the data of the machine and the outside circuit, and (2) from experi-
ments on the machine at no load and deductions from the results obtained.

First let us see how we can draw the diagram from calculated data. We must
calculate the no-load magnetization curve (sometimes called "the saturation
curve ") of the machine, that is, the curve connecting ampere-turns per pole on the

•See article by Dr. S. P. Smith and W. H. Barling, Electrician, October 16, 1914, for
proof that the pointed and flat-topped m.m.f. distributions have the same fundamental sine
wave, and that this may be taken to represent the curve of mean distribution of m.m.f.

ALTERNATING-CURRENT GENERATORS

281

field-magnet with the flux per pole, or the mazimiim flux-density in the air-gap.
Such a curve is given in Fig. 301. The method of obtaining it is given on page 321.
According to our method of calculation, it is most convenient to make the ordinates
represent flux-density in the air-gap. The voltage generated is proportional to
the flux-density in the air-gap. Next, we must draw a curve such as that drawn
with the dotted line in Fig. 301, which shows the increase in the ampere-turns
required to drive the working and leakage flux through the magnetic circuit when

IMAMi

13000

Jft**

0=^

.^/

7IW/

'^^

frtf?

uOOO

-^
^

frow

"Me

Body

fCu

fvc

IIOOO

volts

Yokt

::iy

-ti

ie^

•^ nooo

...

2jgg

^foUs

<i>

'\i>fi

, 0

*
*
*

^ 8000
%7000

^sooo

^ 4000
.2 $000

zooo

lg%J%Jt

lOOO

f Mk

90 ZC

00 JI0

W 46

00 5i

wi a

too m

00 Oi

«V JU

W lOi

W H

00 m

00 I3k

900 /4i

WiSi

WO MH

Ampere Turns per pole

Fio. 801. — ^Magnetization curve of 750 K.y.A. generator.

we have the increased pole leakage that occurs at full load. For instance, in
Fig. 301 at no load the ampere-turns required to drive the full- voltage flux, that is,
to give 9500 c.G.s. lines per sq. cm. in the gap, is 5760. Now, it is shown later
that at full load 0*8 power factor (that being the specified power factor) the extra
ampere-turns absorbed on the field iron due to increased leakage is 800. Draw
the horizontal line NN' to scale to represent 800 ampere-turns from the point N
on the no-load magnetization curve, and thus get the point N^ on the dotted curve.
Thus a new curve can be drawn, giving the increase in the ampere-turns due to
extra leakage on load.

282 DYNAMO-ELECTRIC MACHINERY

We will now suppose that the no-load magnetization curve and the increase-due-
to-leakage curve have been drawn, and that it is required to obtain the relation
between the ampere-turns on the field-magnet and the voltage when full-rated
current is flowing in the armature at a specified power factor, say 0*8. First
calculate the armature ampere-turns per pole from the formula :

0-4:37 1 aZa r i

No.ofpoles=^-P"'P^^^-

Set this off as a vector to scale * as in Fig. 300. The quantities in this figure
are taken from the example worked out on page 321. This vector also represents
the armature current to another scale. Next set off Et = 2100 volts, the terminal
voltage at the correct angle </> according to the power factor of the load. In this case
<^ will be such that cos <^ = 0*8. The armature resistance is usually so small in com-
parison with the self-induction that we may usually for this purpose neglect the
armature ohmic drop, but in Fig. 300 it is shown by the vector lata- For most
commercial calculations it is sufficient to guess at the armature reactance voltage
from a general knowledge of the design. It is usually between 5 and 10 per cent, of
the generated voltage. If it is required to work it out more exactly, this can be done
by the method described on page 388. In the present case the reactance voltage,
8 % of full voltage, is set off by the vector laXa, and thus we get the vector Eg^
representing 2220, the generated voltage. Referring now to the increased leakage
curve in Fig. 301, we find that to generate a voltage of 2220 volts, we require an
excitation of 7700 ampere-turns per pole. Set off the vector I^r to represent 7700
ampere-turns at right angles to Eg. The field ampere-turns per pole are then repre-
sented to scale by the vector I^, which is obtained by subtracting Iza from Izr .

The direction of the resultant field lies along Eg . This is in general not the same
as the mechanical centre-line of the poles shown by the dotted line NS. The
armature ampere-turns Iza operate partly as a cross-magnetizing M.M.F., distorting
the crest of the field form to one side of the centre line, and partly as a demagnetizing
M.M.F., weakening the effect of Izf the applied ampere-turns. The angle ^ between
the current vector and the centre-line of the poles is sometimes spoken of as the
" internal displacement angle." The resolution of Iza into its two components
Ize And Izd is considered later in Fig. 314.

If we replace the curves which represent the distribution of the m.m.f.'s along
the air-gap by their equivalent sine-wave distributions, we should get a diagram
like that shown in Fig. 302, f in which F-^ represents the distribution of m.m.f.
due to /^, and F^ represents the m.m.f. distribution due to the armature ampere-
turns. ^2 ^ shown resolved into two components Fq , the cross-magnetizing effect,
and Fg , the demagnetizing effect. The angle yp shows the displacement of the centre
of the current phase-band behind the mechanical centre-line of the pole.

Suppose now that the machine is run with the armature short circuited, and
that the field current is brought up to such a value that the armature just carries
full-load current. The voltage at the terminals, Eu will be zero, and the voltage

* Oq p. 280 we did not divide by the number of poles. There we took ampere-turns on two
poles both for the armature and field.

t Allgemeine Elektrotechnikf Bd. iii.

ALTERNATING-CURRENT GENERATORS

283

generated wiJl then be Ei^ the voltage required to drive the current through the
impedance of the armature. Some flux, though very little, is required to generate
Ei, Let the resultant ampere-turns reqmred to drive this flux be represented by
the vector /,i . In order to get this resultant, it is of course necessary to more
than overcome the armature ampere-turns /aa- We see, therefore, that the field
ampere-turns required to drive full-load armature current on short circuit is
represented by a vector la, which is greater than /«, by an amount /«• depending
on the value of the armature impedance. If the iron of the machine were not
saturated the extra ampere-turns /»■ required on account of the impedance of
the armature would be the same at all voltages, and we would have a very simple
construction for finding the ampere-turns on the field at fuU load from the ampere-
turns required on short circuit and no-load magnetization curve. We will neglect
the effect of the armature resistance for the sake of simplicity, though it can be

Fio. 302. — ^The sammation of armatuie and field magneto-motive forces.

taken into account in the construction if we wish it. In Fig. 304 set off /« to
represent the ampere-turns on the field-magnet on short circuit. This is of two
parts, Iga the true armature ampere-turns, and /»• the ampere-turns required to
drive the flux which generates the e.m.f. that overcomes the armature impedance.
Let Et represent the terminal voltage, <^ being the angle of lag, and Eg the
generated voltage. At right angles to Et set off 7^ , the ampere-turns required to
give the voltage Et at no load, and set off !„ downwards from the end of Z^^.
Then I^ gives us the ampere-turns at full load. The relation of this figure to
Fig. 300 is seen if we insert the vector Izr- In a word, we have set off Izt at right
angles to Et and added Izs instead of setting off Izr at right angles to- Eg and
adding Iga- The advantage is that I a is obtained at once from the no-load
magnetization curve of the machine, and Izs is obtained from the short-circuit
test. Thus the triangle Igtlzgl;^ is all that is required to find the full-load
ampere-turns, if we can neglect the saturation of the iron and the resistance
of the armature. The correction for the resistance of the armature is obvious

284

DYNAMO-ELECTRIC MACHINERY

from Fig. 300. It merely has the effect of throwing Izi out of line with /«»
(see Fig. 303). The correction for the saturation as given in Fig. 301 is two-
fold. In the first place, Eg being greater than Et will, if the field is saturated,
call for an increase in the exciting current greater than Izr* In the second place,
the ampere-turns on full load cause much more leakage than at no load, and the
leakage flux causing extra saturation again calls for more ampere-turns. For

Fio. 303.— Vector diagram of a
short-circuited generator.

Fig. 305. — Simplified constraction
for finding full-load ampere- tumB.

Fio. 304. — Showing the construction for finding the ampere-
turns per pole for a load of any power factor.

approximate calculations, however, these effects are offcen neglected, and the
simple triangle of Fig. 305 is used in conjunction with the magnetization curve of
the machine in the manner described below.

This then brings us to the method of finding the exciting current at full load
by means of data obtained from measurements made at no load. By running the
machine at normal speed at various excitations, and measuring the voltage generated,
we obtain the no-load characteristic. The curve marked E in Fig. 306 is such
a curve, relating to a 4-pole, 2000 k.v.a. 5000 volt three-phase generator of 50

ALTERNATING-CURRENT GENERATORS

285

periodfl, whose full-load current per phase is 232 amperes. By running the machine
with the armature short-circuited through ampere-meters, and measuring the field
current for various values of the armature current, we obtain the short-circuit
characteristic such as that marked it.

We have seen from Fig. 303 that the field-current on short circuit is made up
of two parts. One part supplies the ampere-turns necessary to overcome the arma-
ture ampere-turns /«», and the other supplies the ampere-turns Igi necessary to
give sufficient flux to generate the voltage that drives the current through the
impedance of the armature. None of the experimental methods that have been

Vofl

/uuu

1 —

•

6000

■

(^A^

5000

4000

3000

2000

}

iOOO

... 1

500

300

200

too

50 . 200 ISO 200 Amp. "

FiQ. 306. — No-load and shoit-ciicuit charaeterittics of a 2000 K.y.A. 5000 volt 8-phase generator.

proposed for separating the short-circuit current into these two components are
satisfactory in practice. It is usual to deal with the short-circuit current in its
entirety as shown in Fig. 305. The ampere-turns required to give full voltage at
no load is represented by a vector ifo as in Fig. 305, and the short-circuit current,
tjt is laid off at the correct angle as shown, where </> is the angle of lag. Then t//
would be the full-load field current, if there were no saturation. The triangle in
Fig. 305 is simply a copy of the triangle Iztlzgl^ in Fig. 304 to a different scale.
For many purposes this simple method is sufficient, especially where our know-
ledge of the machine in question enables us to make a mental correction for the
increase in the current due to saturation. It will be found sometimes that through-
out a whole line of standard machines the ohnxic drop in the armature is about
the same percentage of the normal voltage, and is approximately known. Similarly,

286 DYNAMO-ELECTRIC MACHINERY

the reactance drop in the armature is usually known approximately, and where
this is so the part Izi can be calculated and subtracted from the short-circuit field
ampere-turns, leaving Iza^ from which the field amperes required to overcome the
armature demagnetizing effect can be calculated.

When the amount of the impedance of the armature winding is known, a
skeleton diagram such as that shown on Fig. 307 will be found useful where it
is required to find the field-currents at various power factors. Copies of this
skeleton diagram can be kept at hand, with the radial lines at both ends already
drawn in position. The radial lines on the left-hand side are for setting ofi the
armature ampere-turns, and those on the right for setting off the armature
impedance drop. The method of using this diagram will be best understood by
working out an example.

Suppose that in the machine to which Fig. 306 refers, the reactance voltage
of the armature with full-load current flowing is 8 % of the normal voltage, and
that the ohmic drop in the armature is 1 % of the normal voltage. The horizontal
line OE (which may be conveniently 100 units long) stands for unity or full normal
voltage. The small semi-circles on the right-hand side mark off 5 % and 10 %
reactive drop respectively from the little radial lines. Suppose that we wish to
find the exciting current of the machine in question when operating at full load,
0*8 power factor lagging. We mark off 8 % reactive drop on the little radial line
marked 0-8 power factor lagging, and then set off the ohmic drop tangentially,
and arrive at the point E'%, Then the line OE-^ is proportional to the generated
voltage, = 5300, at the load in question. Referring now to the magnetizing charac-
teristic (Fig. 306), we find the field current required for 5300 volts. This is 109.
Set off 109 amperes along the voltage line to any convenient scale. The field
amperes on short-circuit (armature current 232 amperes) are 52. We can divide
this into two parts just as !„ was divided into two parts, Igi and /«,. The
part required to generate the 400 reactive volts is (from Fig. 206) 7 amperes,
so that the true demagnetizing part is 45 amperes. Set off the 45 amperes to
scale, along the left-hand radial line which corresponds to a power factor of
0-8 lagging. Completing the triangle, we find that the full-load field current is
141 amperes. And so for any other power factor. For cos <^ =0, we see that by the
construction we merely add the 8 % or 400 volts to the 5000 volts, and find from
Fig. 306 the corresponding field current, and then add the 45 amperes directly
to it. This is shown in Fig. 306, and the characteristic for full load cos <^ =0 is there
plotted for various voltages.

To find the field current at half load the same construction is adopted, except
that the impedance drop is taken at half the value and only 22*5 amperes are laid
off along the radial line instead of 45 amperes. And so for any load.

Having obtained the field current at any load, it is easy to find the rise in
voltage which takes place when that load is thrown off, by finding &om the no-
load characteristic the voltage that would be generated by the increased field
current and subtracting from it the normal voltage. If we plot the percentage
of normal voltage obtained when various loads at various power factors are thrown
off, we get curves like those given in Fig. 307a for cos<^ = l and cos <^=0-8, which
have been calculated from the data given in Fig. 306.

ALTEBNATINQ-CUBRENT GENERATOBS

288

DYNAMO-ELECTRIC MACHINERY

It is of interest to enquiie how the field current varies with the power &ctor
when the annatoie is canjing full-load current. It will be seen from Fig. 308

.i._.

_ ._J

fii\i

««

■^

- .1/

t,ts?J!»i

--^

i§*U%

ctif'.

120

^^^

—

— ■

. — '

=—

■

100 ^WO SOOAmp.

Fio. S07a. — Showing rise of voltage with load thrown off.

that for small changes in the power factor the field current changes quite consider-
ably. There is very much more change in the field current in changing the power
factor from unity to 0*95 than in changing it from 0*95 to 0-9 ; as the power factor

F,
tea

f%fV

tAJt

t^V

tzo

#/!/>

too

—

..^

49A

0£ o^ a^ 0-2 o

Fio. 808. — Corve showing relation between exciting current at full load and the power factor

load.

O 02 o^ a€ O^

Bene
of the

gets lower the effect of changing it becomes less, until at cos <^ =0 the rate of change
becomes zero.

The curve showing the relation between the change in the voltage with the
change in the power factor, on a machine having as much saturation as that shown

ALTERNATING-CURRENT GENERATORS

289

in Fig. 306, and having considerable ohmic drop in the armature, differs in shape
from the curve connecting field current with power factor. This will be seen from
Fig. 309, which has been plotted from the data of the machine to which Fig. 306

i€lb

—

r~—

— "■

——

—

T *

100

Zc^

ptn

1 6 a€fir\

0-2 04 o€ oe

OS 06 04- 02

Fio. 800. — Curve showing relation between the change of pressnie when load is thrown off

and the power factor of the load.

refers. It will be noticed that for leading power factors the two curves are nearly
the same shape, but at lagging power factors the effect of the saturation and the
ohmic drop is to make the curve in Fig. 309 almost straight.

Fio. 810. — Field-current diagram for 750 K.y.A. generator.

Strictly speaking, the magnetization curve to be used to find the field current
on load from the generated voltage should be the curve corrected for increase-due-
to-leakage, as shown in Fig. 301. But in practice one commonly makes a mental
correction for this.

In Fig. 310 \a given the skeleton diagram worked out for the power factors,
10, 0-8 and 0-7, from the data relating to the 750 k.v.a. machine calculated

W.M. T

290

DYNAMO-ELECTRIC MACHINERY

on page 321, whose magnetization curve is given in Fig. 301. In this case the
increase-due-to-leakage curve has been employed (see page 330). By means of the
construction given in Fig. 310 it is possible to plot curves which show the change
in voltage which occurs when any load at any power factor is thrown on. Such
a series of curves are given in Fig. 311.

In practice, however, regulation curves are not so much required to show what
any particular machine will do, as to tell the designer what frame he must employ
in order to meet a given regulation guarantee. For this purpose curves of the type

8000

tos^

<IM/l/>

7000

\cosH

€000

si=J=^^

SOOO

""""^

""^

'^--.

•*«-,

^^^

cos

<P'f

1 '

HOOO

cos^

^■p

3000

^N^

Tv^

2000

^■K\

\ 1

k \

1000

;

200

300

WO 500 600 700 Amp.

Fia. 811. — Showing how the pressnie of a 6000-volt generator varies when loads of various
power factors are thrown on. The two upper curves relate to leading power factors.

shown in Fig. 312 are of great service, because they relate not merely to one machine,
but to any polyphase generator. We take for abscissae in this figure the ratio

^ - . ^ '-^. — j, because upon this ratio the regulating quality of the

field A.T. on no load

machine depends. By ^' armature a.t." for any given armature current we mean
(in this connection) the number of ampere-turns that must be put upon the field-
magnet in order to make the given armature current flow through a short-circuited^
armature. The curves are obtained by drawing a number of triangles, such as
shown in Fig. 305 with various ratios between %k and i/o . The abscissae give the

ratio ?- and the ordinates marked *' percentage regulation up '' give the percentage

by which ifj is greater than ijo for various power factors. If, now, we take various
ratios between {r and t//-, and find the percentage by which ifo is less than ijj^ we
find the " percentage regulation down " plotted in Fig. 312 for various power
factors. These curves give the regulation as it would be on an unsaturated
generator. Where we have the no-load magnetization curve of any machine given,
the regulation on that machine can be ascertained by taking the change in field

ALTERNATING-CURRENT GENERATORS

FlO. 812. — Carres for finding rapidly the regulation of any generator at any power factor
or for finding the ratio of armature a.t. to field a.T. for any required regulation.

292 DYNAMO-ELECTRIC MACHINERY

current from Fig. 312, and finding from the magnetization curve the corresponding
change in the voltage.

A few examples will clearly illustrate the method of using these curves.

Example 43. It is required to find the rise in voltage whioh will oocur on the machine
referred to in Fig. 306 when 2000 K. v. a. , at power factor 0*8, is thrown oflF. From the curve ig we
see that a current of 232 amperes in the armature on short circuit requires 52 field amperes.
From curve E, 5000 volts at no load require 96 field amperes. Therefore the ratio armature
A.T. to field A.T., g2

From the abscissa 0*54 in Fig. 312 run up to the 0*8 power factor curve, and we find the
increase in excitation is given as 40%. This would give us 135 field amperes, assuming no
saturation, and according to that the voltage would rise to 5800 volts on the load being thrown
off. The more complete construction given in Fig. 307 gives us 141 amperes for the exciting
current, and the voltage would rise to 5860 volts. For field-magnets of the cylindrical tj'pe,
the use of the curves in Fig. 312 does not give as accurate result-s as the construction given in
Fig. 307. We shall see later that for machines with salient poles the method given in Fig. 312
gives exciting currents which are rather too high. That is to say, the figures obtained are on
the safe side (see page 365, where a case is worked out by three different methods).

Example 44. The 2180 k. v.a. generator referred to on page 348 is running at no load with
an excitation of 96 amperes. A load of 116 amperes (half load) at a power factor of 0*95 i.s
suddenly switched on. How much will the voltage drop, assuming that the speed and excitation
remain constant ?

As before, we find the ratio of armature ampere-turns to field ampere-turns

1=0-275.

From the abscissa 0*275 in Fig. 312 we drop a perpendicular to the 0*95 power-factor our\»e
(regulation down), and we find that the armature reaction weakens the excitation by 11%.
Referring now to page 348, we see that an excitation of 85 amperes gives a voltage of 4700.

The fiixing upon a frame for the building of a generator often depends upon the
maximum number of ampere-turns that the frame can carry without over-heating.
It is therefore necessary that we should be able, when given the maximum field
ampere-turns that a frame will carry, to say what regulation that frame will give
when the armature is loaded at a definite current loading.

The curves in the upper part of Fig. 312 enable us to do this quickly for the
case where cos<^=0-8, and with suflScient accuracy for the purpose of choosing
the size of frame. An example will make the matter clear.

Example 45. Suppose that we have a 16-pole frame that will carry a maximum of 10,000
ampere-turns per pole, or 160,000 ampere-turns total. If we put a current loading on this
frame of 124,000 ampere-wires, what kind of regulation would we expect to get at 0*8 power
factor ? The ampere- turns on the field to give full-load current on short circuit, I„ (see
Fig. 303), may be taken as very nearly one-half the ampere wires, in this case 62,000 ampere-
turns.

n, , .V ^. field A.T. on full load 160,000 ,, ^

Take the ratio =-7r., ^7^=2-6.

armature a.t. 62,000

Take the ordinate 2*6 on the right-hand side of Fig. 312. Run along the horizontal line
until we come to the curve marked "no increase." Then drop vertically to the abscissa scale
and read off 0*54. This is the ratio of armature a.t. to field a.t. at no load. From this we can
run up to the curve marked power factor 0*8, and find that the field current will have to be
increased 40 % on that power factor. The actual regulation will depend upon the amount
of saturation in the magnetic circuit (see Example, page 293).

ALTERNATING-CURRENT GENERATORS 293

Now, the curve marked " no increase " has been plotted on the assumption
that there has been no increase in the field current at full load due to saturation.
As this assumption is not warranted in machines in which we are relying upon the
saturation to improve the regulation, the two curves marked 12^ % and 25 %
respectively have been introduced. The first of these is plotted on the assumption
that the saturation of the frame is such that when the ampere-turns on the armature
are made equal to the no-load ampere-turns on the field, there is an increase in the
field ampere-turns on a load of 0-8 power factor of 12^ % over the amount cal-
culated by Fig. 305 due to the saturation of the magnetic circuit. For smaller
current loadings a smaller allowance is made for saturation. This curve is the one
which will generally be used with a.c. generators as commonly constructed. The
25 % curve and other curves which we can imagine to be drawn in between the
two are intended to be used when we are dealing with more highly saturated
machines.

Example 46. A certain frame cjan carry a maximum field a.t. of 175,000. If a generator
built upon it is designed so that the ampere- wires /o-^a= 120,000, and if the amount of saturation
in the magnetic circuit is the usual amount as indicated in the characteristic curves in Fig. 306,
what voltage rise will we get when full load at 0*8 power factor is thrown off?

As before, take the ratio ^^!^^^=2-9.

120,000

Running along the line 2*9 until we come to the 12 J % curve and then downwards, we get
the abscissa 0*47 for the ratio between armature a.t. and no-load field a.t. At this ratio the
increase in the field current at full load 0 8 power factor is 33 %. Referring now to the no-load
characteristic, we find that an increase in the field current by 33 % will give a rise in the voltage
of 14 %.

These curves are, of course, only intended for obtaining approximate results.
By means of them we can save a great deal of time in preliminary calculations.
We can, for instance, find out in the course of a few minutes whether it is possible
to squeeze a generator of a certain output upon a certain frame and yet stand a
good chance of meeting certain guarantees as to regulation. They at the same time
tell us, for any particular frame, what ratio we may take between armature a.t.
and no-load field a.t., and yet not exceed at full load the maximum a.t. of the
field that we know must not be exceeded.

The full-load ampere-turns for the purposes of the ratio set out on the right-
hand side of Fig. 312 are taken as the full-load a.t. at 0-8 power factor, that being
the power factor for which machines are most commonly designed. It would be
necessary to plot other curves if the ratio between armature a.t. and full-load
field current at other power factors were the basis of the calculation.

The method given above for the calculation of the exciting current of a loaded
generator is strictly only applicable to a machine with a cylindrical field-magnet,
such as is illustrated in Fig. 371, in which the reluctance of the magnetic path is
the same in every direction. Where a machine has salient poles, the reluctance
of the path for the cross-magnetizing flux is higher than for the working flux, and
therefore the vectors representing fluxes are not proportional to vectors representing
magnetomotive forces. The circumstance is only of importance where it is neces-
sary to calculate the regulations at high-power factors as accurately as possible.

294

DYNAMO-ELECTRIC MACHINERY

For ordinary commercial purposes the method given on page 284 is commonly
applied to salient pole machines, because it gives results which are sufficiently
near for practical purposes, and the error that exists is on the safe side.

The more exact method of calculating the exciting current will be understood
from the example worked out below. It differs from the method given on page
284 mainly in the fact that it is necessary to resolve the armature magnetizing
effect into two parts, one acting directly against the field ampere-turns and the
other acting at right angles to it and constituting a cross-magnetizing effect. This
latter magnetomotive force must be midtiplied by a coefficient, depending on the
ratio of pole arc to pole pitch, before we can arrive at the effect it will have in

(TV

ftmM

fr9

»S A.A.

0-1

0 •/ -Z -3 ■-# -5 -€ '7 -9 -9 t-0

Ratio Po^^
Pole pitch.

Fio. 818.

shifting the phase of the e.m.f. behind the centre line of the pole. This coefficient
we denote by Kq . Fig. 313 shows how the coefficient Kq changes with the pole
arc on ordinary three-phase generators having poles of the kind shown in Fig. 234.

The method is one of trial and error, because the angle \^ between the phase
of the current and the centre line of the pole, " the internal displacement angle,"
depends upon the amount of the cross-magnetizing effect, and this again depends
upon ^.

A diagrammatic view of a generator having two salient poles is shown in Fig.
314. The arrangement of the armature conductors is the same as in Fig. 300,
and we may, for the purpose of our example, take the armature ampere-turns on
the full-load and no-load characteristic the same as in the example worked out on
page 282.

The method of finding the field ampere-turns on full load is as follows :

Lay off the vector Ixa to represent the full-load armature ampere-turns. This
vector will point to the centre of the phase band A when the phases of the currents
are as indicated in the figure. Lay off the vector Et to represent the terminal

ALTERNATING-CURRENT GENERATORS

295

X.M.F., the angle <^ being the angle of lag between current and voltage according
to the power factor of the load. Then, as before, lay off lata and laXa and so obtain
the generated voltage Eg, We know that the phase of this generated voltage
lags behind the phase of the centre line of the pole by an amount which depends
npon the cross-magnetizing effect of the armature, and as a consequence the current
lags further behind the centre line of the pole than it otherwise would. We deter-
mine the value of the angle ^ between the centre line of the pole and the current
by a method of trial and error.

Fig. 814. — Showing construction for finding the excitation on full load of a salient pole

d-pbase generator.

First, assume that ^ is the same as it is in Fig. 300 and resolve Im into two
components, one, 7^., along the supposed centre line of the pole, and the other,
Izdi at right angles to it. The component Izd is a direct demagnetizing force, when
the current lags behind the centre line, and the component Izc represents the cross-
magnetizing magnetomotive force. Each of these vectors is drawn in the direction
in which the corresponding current components would flow, so that the directions
of the corresponding magnetizing forces are, of course, at right angles to these
vectors respectively. Now, the reluctance of the magnetic path through the field-
magnet at right angles to /»; is greater than the reluctance of the path at right
angles to ltd • Hence the necessity of introducing the coefficient Kq . Scale off the

296

DYNAMO-ELECTRIC MACHINERY

provisional Izc (shown dotted) and multiply by Kq for the ratio of pole arc to pole
pitch in question. In this case the pole arc is 0-65 of the pole pitch, so from Fig. 313
Kq =0-32. To the same scale as the abscissae of Fig. 301 the provisional Izc represents
2000 ampere-turns. This multiplied by 0-32 is 640 ampere-turns. From Fig. 301
this would generate a cross voltage of 333. If we set off this voltage in the position
shown at Ec, we will find that the provisional line taken for the centre line of the
pole was not right, but the true position is very nearly indicated. A second trial
gives us the true position as marked in Fig. 314. We can now resolve Im into its
two components, /«• and Izd» much more accurately. Ize now represents 2300
ampere-turns ; multiplying this by 0-32, we get 736 ampere-turns, which yield
380 volts according to Fig. 301. Setting off Ec to represent 380 volts at right angles
to the centre line of the pole, we get E.^,, the e.m.f. generated by the undistorted
flux from the pole.

Fio. 315. — Diagram of the cross maffnetomotive force Fq and the croBs flux Nq on a salient

pole S-phase generator.

In Fig. 314 we have shown a two-pole machine, so that the meaning of the
vectors can be made more apparent. The same construction is applicable to
machines having a large number of poles. In Fig. 315 the large dots represent the
centre points of the phase band of the component of the armature current, which is
in phase with the centre line of the pole (compare Fig. 302). This band of current
lying in a distributed winding will yield a magnetomotive distribution represented
by the wave-form Fq, Taking into account the reluctance in the various parts
of the magnetic circuit, we see that this distribution of magnetomotive force would
produce a flux distribution, something like that indicated by the thicker curve
Nq. This flux distribution, when combined with the no-load flux distribution shown
at No in Fig. 316, gives the resultant flux distribution shown by the curve N,

It will be seen that this cross-flux distribution, if acting alone, would generate an
E.M.F. (denoted here by j&c), which would have a very pronounced third harmonic,
but if the armature of the generator is star connected this harmonic is entirely
neutralized, and the Ec that remains is (with a pole arc 0-66 of the pole pitch)

ALTERNATING-CURRENT GENERATORS

297

only about one-third of what it would be if the air-gap were of the same length
over the whole pole pitch (see page 308).

The no load field-form for a salient pole is shown by the curve Nq in Fig. 316.
If the armature magnetomotive distribution is as indicated by the curve F^t the
resultant field-form might be somewhat as indicated by the curve N, The exact
form of this curve depends upon the amount of bevel on the poles and the state of
saturation of the iron, so that without going into each case with great minuteness
it is impossible to say exactly how much the field is distorted. The values given
in Fig. 313 for the coefficient Kq are sufficiently near for practical purposes where
the poles are of the general shape shown in Fig. 234.

For ordinary standard machines with a pole arc of about two-thirds of the
pole pitch we are justified in applying the vector diagram as described in Fig. 305,
notwithstanding the fact that the cross-magnetization is not wholly operative.

Fig. 816. — ^Resultant, N of main flux No and cross flux Nq (Fig. 815).

It will be seen that the regulation of the machine will depend both on the ratio
of the field ampere-turns to the armature ampere-turns, and on the extent of the
saturation of the field system. In designing a machine to comply with any par-
ticular regulation guarantee, we may either make the ratio of field ampere-turna
to the armature ampere-turns sufficiently great and have little saturation, or we
may make the ratio less and have more saturation. If we wish for a generator
with large overload capacity, we will adopt the first course. If we wish for a cheap
machine with smaller overload capacity, we will adopt the second course, and we
may adopt any intermediate course according to circumstances. If the ratio of
field ampere-turns to armature ampere-turns is made too small, and an attempt
is made to secure the regulation by excessive saturation, there is danger that at
low-power factors it will be impossible to obtain the specified voltage at all.

It should be pointed out that the effect of the saturation in limiting the output
of a machine greatly depends on the part of the magnetic circuit in which the
saturation takes place. If it occurs in the teeth of the armature only^ then it would
not prevent us from obtaining full voltage at heavy loads. It is true that a little
more flux is required at heavy loads, because the generated e.m.f. is greater, but
even with high saturation in the armature teeth there is not a call for an excessive

298 DYNAMO-ELECTRIC MACHINERY

number of ampere-turns to make a small increase in the total flux per pole. If,
on the other hand, the great saturation occurs at the root of a rather long pole,
or worse still, in the yoke itself, then it may easily happen that we cannot get even
the no-load flux through the armature at heavy loads of low-power factor. The
armature back ampere-turns call for more ampere-turns in the field, and these may
increase the leakage to such an extent as to rob the armature of some of its flux.
A further increase in the field ampere-turns causes a further increase in the leakage,
and so on.

A good plan is to allow the saturation to occur on the surface of the pole.
Where the mechanical construction permits of it (as, for instance, in cylindrical
turbo-rotors) saturation may be produced by cutting slots near the surface of the
pole. This gives a much more satisfactory moignetization characteristic than can
be obtained from a simple salient pole saturated at the root. Figs. 352 and 353
show the difference between the no-load and the full-load characteristic in
the two cases, the reason for the differences being as stated above. Where a
punched salient pole is employed in ordinary engine-type machines, it is possible
to punch slots near the pole face. This, however, requires a special die, and
would not be economical unless a large number of machines (or rather a
large weight of punchings) were required, using the same die. A somewhat
better plan (where comparatively few machines are to be built) is that shown
in Fig. 334. Here the saturation occurs just below the pole cap. This is not the
best place, but it is not far removed from the best place. This plan has the advan-
tage that the space saved by the cutting out of the iron can be utilized as copper
space. The best length of pole to saturate depends on the pole pitch. With a
pole pitch of 12', a saturated length of 3" gives good results. The extent of the
saturation is a matter which requires very careful adjustment. It will be seen from
the cases worked out below the sort of considerations that determine the best
amount of saturation.

Before we can fix upon the size of frame upon which to build our A.C. generator,
we must not only know the regulation required, but we must decide whether we
are going to obtain that regulation by giving a sufficient ratio of field ampere-
turns to armature ampere-turns, and relying very little on saturation, thus obtaining
a machine of great overload capacity ; or whether we will rely to a great extent
on saturation, and build a cheaper machine, sufficiently ample to do the work
it is designed for.

If we adopt the first plan, the curves in Fig. 312 are useful in arriving at ampere-
turns required at full load, and in the choice of the frame upon which to begin the
design. Suppose it is specified that the voltage shall not rise more than 25 %
when full load at 0*8 power factor is thrown off. From Fig. 312 we see that if the

ratio

field amperes at short circuit _ ^.

field amperes at no load '

we will require an increase of about 28 % in the field current when on load. Now,
with only the very smallest saturation, an increase of field current of 28 % will not
give more than 25 % rise. So we may take the ratio 0*4 as sufficient to meet
the guarantee.

ALTERNATING-CURRENT GENERATORS 299

Now the three-phase output of a frame in k.v.a. is equal to

Ke X Rpm, X AgBk X 10-» X ZoZa X 1-73. (See page 24.)

ZoZa is limited by the ampere-turns which the field-magnet is able to carry
at full load without exceeding the guaranteed temperature ftSe? The total armature
ampere-turns, as we have seen (page 280), are equal to 0437 Ic^at and the short-
circuit field ampere-turns may be taken roughly at 047 laZa . Divide this by 04,
and we get the field ampere-turns at no load ; multiply by 1 -28, and we get the
field ampere-turns at fuU load. Thus we have, in this case, l-5/aZa= field ampere-
turns on full load 0-8 power factor. Let the factor by which we multiply the /^Za
to get the field ampere-turns on load be denoted by Kr . The value of Kr will depend
upon the regulation required.

Then KrIaZa = IfS x 2p, where 8 is the number of turns per pole and 2p equals
the number of poles, and // is the field current. If now we have all our frames
tabulated so that we can tell at a glance what magnetic loading, AgB, and what
maximum number of ampere-turns, 2pI/S, each frame will take, it is a simple matter
to fix on a frame. When we have not these data available, it ia necessary to employ
^ JDH formula to give us the approximate size of frame. Now as the ordinary
BH formula takes no account of regulation, it is a good plan to modify it in the way
given below. It will then be a useful guide in the choice of a frame where the
regulation is specified. A 16-pole 50-cycle a.g. generator with a peripheral speed
of 6000 feet per min., and with the iron and copper space well adjusted, will carry
about 10,000 ampere-turns per pole for 45° C. rise. As the pole pitch is 12 inches,
this amounts to 850 ampere-turns per inch of periphery. So that if we multiply
vD" by 850, we get the possible number of ampere-turns on the field of diameter
Bf. The ampere-turns per inch of periphery depend largely on the pitch of the
poles, and on the kind of winding, as well as on the peripheral speed. Fig. 318
shows how the economical number of ampere-turns per pole changes with the
•diameter of the machine and the number of poles. We shall return to this matter
later ; for the moment we are modifying the BH formula to provide for the regu-
lating qualities of the machine, and we assume that we know the figure 850 for the
<;ase under consideration.

We have the total ampere wires, /aZa = -•^-~ — ^= = .

If we take^^Sfc = DVJ x 10 (see p. 6), then

output in K.V.A. = JSTe X Rpni X AgBic X 10"« x IJ^a X 1 '73

= ^ xRp„,xD^lx7r^x 850 X lO"® x 1 73.
Taking Kg at 0 4, this becomes

K.V.A. = pT X Z>*? X fip„» X 850 X 6-85 x lO"®.

Thud we have the K.V.A. in terms of the diameter and length (in inches), and
the regulation constant Kr* It remains to consider how the field ampere-turns
per iAch of periphery, which we have here taken as 850, change with the size of
the machine and the number of poles. In Fig. 317 we have drawn the poles and
<;oil8 in 3 cases : (1) the 8-pole case, (2) the 16-pole case and (3) the 32-pole

300

DYNAMO-ELECTRIC MACHraERY

case, for a rotating field-magnet 60" in diameter. The length of the iron azially
is supposed to be 12J inches. As explained on page 277, the ratio of the width of
the pole to the pole pitch is chosen from certain economical considerations.
Where the poles ar^igw, as in the 8-pole ease, the mechanical support of the
coil is also an important consideration. The ratio of the length of the pole to
the pole pitch is settled by somewhat similar considerations. Where the number
of poles is great, the leakage between the flanks becomes very great if the pole
is made too long, so that the extra copper space gained is somewhat counter-
balanced by the extra ampere-turns required to overcome the reluctance of the
saturated pole base. The proportions shown in Fig. 317 may be taken as
representing good economical practice.

Fia. 317. — Showing arrenBamenUof copper and Iron In mac
dlSenni nambecs of poles.

It must be understood that it is possible to get somewhat more ampere-turns
on the poles than are here considered by adopting special methods of putting on
more copper, and by making ventilating ducts on the ends and sides of the coils.
As the coils are subjected to great centrifugal forces, it is doubtful whether such
devices are altogether to be recommended, particularly as we can quite easily
increase the output of the frame by increasing its diameter. We will consider
first coils wound with square double cotton-covered wire of a size suitable for
excitation at 110 volts. The coils are supposed to be treated with heat-conduct-
ing enamel between layers. We have taken the case of wire-wound coils, because
in general it is more difficult to treat from the heating point of view than the case
of strap-wound field coils. Afterwards we will consider what modifications to make
in our conclusions if strap-wound field coils are employed.

In the 16-pole case we have 118 turns of d.c.c. sq. wire 0'252 inch bare, 0'275
inch insulated. Adopting the method of calculation given on page 233, we will
find that for 40° C. rise by thermometer each coil can dissipate about 554 watts.
Of this, 2ii watts pass through the insulation to the pole and cheeks, 250 are given

ALTERNATING-CURRENT GENERATORS

301

off by the ends of the coil as defined by Fig. 234, and only 60 watts by the sides,
although these have an area of 2(12 x 7) sq. inches.

As the coil has a resistance when hot of '069 ohm, calculation gives us 90 amperes
as the limiting current for 40 degrees rise. 90 x 118 = 10,600 ampere-turns per pole.
We will deduct 10 % from the calculated capacity, and say that we can rate the
•coil at 9600 ampere-turns maximum. From an actual experiment on a coil of this
construction, the actual figures were as follows : Resistance of coil hot '066. Run
for 6 hours at 85 amperes. Temperature rise 34° C.

Now take the 8-pole case. This has 212 turns of the same size of wire as in the
last case. Here the coil has a depth of winding of 2^ inches, so that the tempera-
ture of the surface when running will be somewhat cooler than the interior. This
<;ircumstance makes the rate of cooling per sq. in. of surface of these deep coils
rather smaller than for coils of fewer layers of wire. The application of the formula
given on page 233 to the coils in Fig. 317 in a method of trial and error tells us
that for the 8-pole case the mean temperature of the coil is about 8° C. above the
temperature of the outside, the centre being 13° C. higher than the outside. In
the 16-pole case the difference is only 3° C, and in the 32-pole case less than 1° C.

We must therefore, to make a fair comparison, take the cooling of the surface
in the 8-pole as if the temperature rise were only 32° C. On this basis we will find
that the big coils will only dissipate about 900 watts each for 40° C. rise, and as the
hot resistance is 0*167 ohm, 73 amperes therefore appears to be about the maxi-
mum field current. Allowing again a margin for safety, we may take 14,000
ampere-turns per pole as the safe rating of the 8-pole case.

A similar investigation shows that the 32-pole case can carry 95 amperes, giving
5000 ampere-turns per pole. It should be pointed out that the coil in the 32-pole
-case would be much better made of edgewise- wound copper strap. This would give
a safe rating about 20 % higher, but for the sake of a fair comparison we have kept
to square D.c.c. wire all through.

Table of Data of Revolving Field Magnets.
Diameter, 60*; length, 124''; speed, 375 R.P.M.

No. of poles

Turns per pole
Mean length of turn
Size of wire -
Resistauoe hot
Exciting current -
Amps, per sq. in. -
Amp. -turns per pole
Total amp. -turns -
Amp. -turns per in. perimeter
Weight of copper -

212

118

6(r

45i"

i^"

0-252* sq.

1 0-252 sq.

0-252 sq.

0167

0069

0024

66 amps.

81 '5 amps.

95 amps.

1065

J 1310

1530

14,000

9600

5000

112,000

154,000

160,000

595

815

850

2040 lbs.

1730 lbs.

1200 lbs.

The data are collected in the table above. From this table we see the great
economy of material when the number of poles is increased. Though the W is

302

DYNAMO-ELECTRIC MACHINERY

the same for all macliines and the speed the same, the 32-pole machine can cany
40 % more ampere-turns on the frame, notwithstanding that the weight of copper
in the field-magnet is 40 % less than in the 8-pole case. We therefore cannot
intelligently use any D^l formulae or curves for finding the output of alternators^
unless we take into account not only the regulating qualities of the generator, but
also the effect which the frequency and speed will have upon the number of poles
and the cooling conditions of the field coils.

4
/.

TojDOO

ZSjOOO

/4/k>/ss

2^,000

23t.000

/ePoies

72,000

21 OOO

r
/

/8Po/es

2Q.OO0

9M/1a/a»

t9,000

jaffvfSS

% /6,000

^ ,

22MOS

jJ /7.000

24M9S

(^ 16,000

26fiO/0S

^iS.OOO
\l4.000

^a.ooo

23fio/es

30P0/es .
32Mes

1 a.ooo

y ti.ooo

}

^/QOOi/

^ S,0{M?

:<:;

::^

0,000

7,000

V.'

€/X)0

'//

:<:

0,000

'/y

^OuO

//..

P*^

3,000

»*T*

'*;^

'■>x

'^A

^OOO

/.OOO

too

Inches dUxjneter of Revolving Field-magnet

^ Fig. 818.

In Fig. 318 we have taken as abscissae the diameter of the field-magnet, and as
ordinates the ampere-turns per pole. For ^0" diameter we have obtained three
points, namely, for the 8-, 16-, and 32-pole cases. We have worked out similar
cases for smaller diameters and larger diameters, so as to be able to fill in the curves
as shown. The figures must not be taken as the maximum possible (see page 277)^
but may be taken as good economical values for the full-load ampere-turns per
pole, where the coils are wound with square D.c.C wire of a size suitable for 110

ALTERNATING-CURRENT GENERATORS

303

volts excitation. If copper strap edgewise- wound is employed, and the same weight
used, the ratings can be increased 15 % or 20 %, but as a rule with strap field coils
one saves the 15 % or 20 % of copper and keeps the rating as before. It should
be pointed out that all these cases are for machine 12^ inches axial length with
natural ventilation. For narrower machines the ampere-turns per pole can be
slightly increased. In the 16-pole case the narrowing of the frame to 7^ inches
will increase the possible ampere-turns per pole about 7 %, while for the 32-pole
case the increase would be about 11 %. The 12 J inch length is an economical one
for 50-cycle generators to be driven by high-speed engines. A widening of the
frame will reduce the possible ampere-turns per pole.

«•

800

90O

XfOO

zvs.

pel

"minute

Fio. 319. — Showing peroentago Increaae in the poesible number of ampere-turns per pole as
the speed of an A.O. generator is increased beyond 375 R.P.1C.

Fig. 318 has been worked out for a speed of 375 R.P.M. This speed will not be
suitable for some of the sizes given, so it is necessary to correct the rating for the
change of speed.

We can, on the basis of the rules given on page 233, arrive at the eflEect of the
change of speed on the possible number of ampere-turns per pole. The matter is
complicated by the fact that for different frequencies the economical depth of
copper is different. For machines between 25 cycles and 60 cycles, however, we
may use Figs. 318 and 319 and arrive at a fair idea of the possible number of
ampere-turns per pole for a field magnet of given diameter. Thus, at 60" diameter
with 16 poles, we would get a 50-cycle machine running at 375 r.p.m. with 9600
ampere-turns per pole at full load.

If this frame were used for 25 cycles running at 187'5 r.p.m. we could with
the same temperature rise have 21 % less than this, or 7600 ampere-turns per
pole. Or if the frame were used for 60 cycles running at 450 r.p.m. we could
have 6J % more ampere-turns, or 10,200 per pole.

304

DYNAMO-ELECTRIC MACHINERY

If we confine our attention to 50-cycle generators, the range of peripheral
speed will, for engine-type machines, be small, so that the range of ampere-turns
per pole will also be small. Fig. 320, which has been deduced from Figs. 318 and
319, is useful for reference in designing 50-cycle generators. The number of poles

10 20

N?ofpoUs

30 40 so 60 70 SO SO 100 UO

DLauneter irv inchjes

I I 1 I I I 1 I l_J \ ! I

S to 12 14 16 18 20 22 24 26 2S 30 32
Fig. 320. — GlTing the possible ampere-turns per pole for 50-cycle generators of different diameters.

given on the lower scale is the number which will be found most commonly used
for frames of the diameter given. The curve giving ampere-turns per inch of
periphery is also useful in finding the possible output of a frame.

THE WAVEFORM OF THE ELECTROMOTIVE FORCE.

Where a generator is provided with open slots on the armature or on the field-
magnet, these sometimes produce ripples in the wave-form of the e.m.f. It is
important that a designer should be able to say in what cases ripples will be pro-
duced, and he should be able to calculate approximately the size and frequency
of the ripples.

If the field-form of the magnet is a simple sine wave without higher harmonics,
and if it moves forward at a constant velocity, the e.m.f. generated in each con-
ductor must be a simple sine wave ; and however many of these are connected in
series, and whatever the difEerence of phases may be, the resultant must be a simple
sine wave.

Where the field-form contains higher harmonics (see page 22), the occurrence

ALTERNATING-CURRENT GENERATORS

305

of these harmonics in the resultant e.m.f. depends upon a number of factors which
are considered below.

Slots and projections on the field-magnet, such as pole-tips, may be regarded
as the origin of the harmonics in the field-form. Slots on the armature may or
may not give occasion for these harmonics to be impressed on the e.m.f. wave-form.

If a slot is skewed (see Fig. 533) by a whole slot pitch, its position is so distri-
buted over the whole slot pitch that a winding lying in it may be regarded as a
perfectly distributed winding, and the slot and tooth effect is completely eliminated.
Similarly, where a machine has many poles, and the slots under one pole take up
positions different from those of the slots under another pole (as in the case where
there is a fractional number of slots per pole), the effect upon conductors lying in
the slots and connected in series is the same as if there were a greater number of
slots under the same pole taking up all the positions found under all the poles.
Thus, with comparatively few slots per pole, we may get the effect of a winding
distributed in a very large number of slots, and the resultant e.m.f. will be free
from spacing ripples * (see Class B, pages 102 and 109).

Where, as is often found in practice, there are comparatively few slots per pole,
and each pole occupies the same position with respect to the slots under it, the
. ripples in the e.m.f. wave-form produced by such an arrangement may be of con-
siderable importance.

Dr. S. P. Smith and Mr. R. S. H. Boulding have made a very complete study of
the ripples occurring in the wave-form of alternating-current machines, and they
have kindly permitted the author to give the following abstract of a paper* which
they have read before the Institution of Electrical Engineers.

. A statement of the subject to be sufficiently comprehensive necessarily involves
an introduction of the winding factors (see page 33) belonging to the various
harmonics in the field-form.

There are two ways in which the flux embraced by a coil may vary — either the coil may
move with respect to a steady flux, or the flux may change in amount with respect to a stationary

Fig. 821. — Showing a phase-band of oonductors of width S moving in magnetic field.

coil. The general case occurs when these two modes of variation happen together. When
a coil (Fig. 321) moves relatively to the flux, an e.m.f. of motion or roktiion is induced in it ;

^ • II.

* Joum. Inst, Elec, Engnrs,, vol. 53, p. 205, 1915.

306 DYNAMO-ELECTRIC MACHINERY

whilst if the flax varies, an E.M.F. of pulaation is induced. Expressed mathematically, these
two ways in which the flux 0 may vary are given by :

where the first term on the right-hand side denotes the change of interlinkages due to motion
and the second due to pulsation.

In a heteropolar machine, the flux interlinking a coil at any instant is (see Fig. 321) : -

= f^Bldx,

where {= core-length over which the flux-density B extends. Hence the instantaneous pressure
induced in Tc turns moving at a constant velocity v=dx/di cm. per sec. with respect to the
flux <f> will be :

- - r.gio-= - r. io-s(^ . J..^) (i>

._^,.10-s|.|/>d.+|/>ei.}

= -r..r.Z10-/;g^-T..M0-/;fcto

= -Tc.v.l(Bjr-B^)10-^-Tc.llO'^r^dxYo\ts, (2>

where / t^ dx= j dB.

With continuous-current excitation, the flux is steady, except in so far as
pulsations are set up by the teeth or by armature reaction. We are here considering
only the effects at no load. The effect of flux pulsations due to the t«eth are con-
sidered on page 313.

When the flux is steady, that is, constant in value and fixed with respect to the

BB
poles, the flux-density B at any point in the air-gap is constant, so that -^ =0,

and the e.m.f. is induced by rotation alone. We then have

e=Tc.v.l{Bx-B^')10'^Yo\U (3)

If the coil spans a full pole-pitch, Ba.= - B/, and Eq. 3 becomes :

6 = 2 .Tc.vA.B^. 10-« volts (3a)

Since this is an instantaneous value, the curve of e.m.f. induced in a full-pitch
coil is identical in shape and phase with that of the flux distribution. Further,
as e changes its sign with B, it alternates with the frequency n cycles per second.

We have seen on page 33 that the instantaneous value of the sum of the e.m.f.,
generated in the conductors a, &, c to m of a uniformly distributed winding, is given
by the expression

'e = 2Tvm-»iB,^'^sine+B,^^sbx3e+...Bj'^^Bmh0\ (4)

I ' o- "So- fur J

where o- is the angle subtended by half the coil breadth = - -!.. The factor — ,

® -^ T 2 ho-

is the " winding factor " for the particular harmonic in question, where A is the order
of the harmonic. Smith and Boulding have given the following table of winding
factors for the odd harmonics up to the 25th for different spans of armature coil
expressed as fractions of the pole-pitch.

^fW

ALTERNATING-CURRENT GENERATORS

307

UNIFORMLY DISTRIBUTED WINDINGS.
Table XV. Values of Winding Factors.

Spread of \ _
Winding S/t/-

1/6

1/8

1/2

2/8

Open winding.
Closed ,,

12-PH.

8-PH.

6-PH.

2-PH.
4-PH.

(1-PH.)
8-PH.

Diam. taps.

0-988

0-955

0-900

0-827

0-637

0-900

0-636

0-300

0000

- 0-212

0*738

0191

-0-180

-0165

0-127

0-527

-0136

0-129

0-118

-0-091

0-300

-0-212

0-100

0-000

0-071

0-090

- 0087

0-082

-0-075

-0058

/l8

- 0-076

0-073

-0-069

0064

0049

/»

-0-180

0127

-0-060

0-000

-0-042

/l7

-0-217

0056

0053

-0049

0037

-0-194

-0050

0-046

0-043

- 0033

-0-129

-0091

-0-043

0-000

0030

-0O43

-0041

-0039

- 0036

-0-028

0040

0038

0036

0033

0025

By means of these winding factors, we can calculate the pressure 2a e induced
in a winding, distributed uniformly over any fraction of the pole-pitch, by any
flux whose wave-shape is known. The above figures are of great interest, for
they show the amount by which the fluz harmonics are reduced in the pressure
curve by the spread of the winding. For example, in a section of winding spread
over the whole pole pitch (Sfr = 1), the magnitude of the winding factor in per cent,
is 100/A, Le, in this case the winding factors bear the same ratio to one another
numerically as the coefficients of the harmonics of a rectangle (see page 22).

The winding factor for any harmonic. A, can also be represented graphically
as the ratio of the length of the chord to the length of the arc in subtending an

angle A-tt radians at the centre of a circle (see page 112). The arc of the circle

represents the actual e.m.f.'s induced in the several coils,, whibt the chord repre-
sents the resultant of these e.m.f.'s, which are slightly out of phase with one another.
Another interesting feature arising from the spread of the winding is the fact
that under certain circumstances harmonics which may be present in the flux curve,
and therefore in the e.m.f. of each conductor also, disappear entirely from the
phase pressure. The conditions for this can be directly deduced from the general
expression for the winding factor. In order that any particular harmonic, h,

shall not reappear in the pressure, 2^6, if present in the B-curve, it is sufficient

and necessary that/^=0, or sinA-^ = 0 (since the denominator h~^ obviously
can never be zero). Now,

sin A - o "=^^111 Xtt =0 only hol^s when X=s — =0, 1. 2, etc

A, of course, being any odd integer.

308

DYNAMO-ELECTRIC MACHINERY

A 2 h
For example, let S/t=2/Z, then X = ^ ^=o* It is at once Been that X will be integral

when, and only when, A =3, 9..., etc. So that with i9/r= 2/3, no harmonic whose order is a
multiple of 3 will appear in the pressure wave. This is an important result, for it shows that
there can never be a third harmonic in the line pressure of a star-connected, three-phase gene-
rator, nor in the alternating pressure of a three-phase rotary converter, nor in the phase pressure
when each phase extends over 2/3 of r, nor in the pressure of a single-phase alternator with
two-thirds of the periphery wound. It will be seen that this also holds when the winding is
placed in slots instead of being uniformly distributed. The absence of the third, ninth, etc.,
harmonics is noticed in the above table in the column where 8/t=2/3.

Again, in a similar way with ^/r=s2/5 or 4/5, it can be shown that no harmonic which is a
multiple of 5 can appear in the pressure curve. The important case in practice, however, is
the one previously referred to when S/t=2/3 ; and it is to be noticed that by making the phase-
band two-thirds of the pole-pitch, it is possible to have a star-connected, three-phase winding
without a third harmonic in the wave-form of the e.m.f. generated in one leg of the star.

In order to arrive at the wave-form of the e.m.f., it is necessary to know the
form of the B-curve, or, in other words, to know the values of the coefficients Bj,
Bj, Bj, etc. (see page 22). Where the B-curve is irregular in form the coefficients
can be determined by any of the methods of harmonic analysis.

It may be that in a machine with a salient pole the rectangular form depicted in Fig. 322
is sufficiently near the truth for the purpose of arriving at the approximate value of the
coefficients. We then have

B =-B^(co6asin9x+icos3asin39x+...).

(5)

The coefficients of the various terms here depend upon the ratio of pole -arc to pole -pitch.

-jr-cr-

FlO. 322.— Rectangular field-form.

FIG. 823. — ^Trapesium field-form.

Taking the ratio of pole-arc, h, to pole-pitch, r, as two-thirds, which is usual for slow-speed alter-
nators, the equation for the flux distribution is found by substituting : a= ^=30° in Eq. 5 :

B* = -B-^(sin ^x - i sin 6dx - f sin 7^,+^ sin 11^,+ ...).

It is seen that with this ratio of pole-arc to pole-pitch, all harmonics whose orders are multiples
of 3 vanish in the flux curve, so that there can be no third, ninth, etc., harmonics in the pressure
waves. For the remaining flux harmonics :

also

and

»=-!. §r=-l. Bn=^l ,tc
o _4„\/3_3\^ 0

ALTERNATING-CURRENT GENERATORS 309

Subetituting these values in Eq. (4)| p. 906, "we get for the constant term :

80 that the equation for the pressure 2^e with a rectangular flux distribution over t-wo-thirds
of the pole-pitch becomes

s;;*e= -^^rn^l0-»(/i8in^-i/j8in6tf-f/78in7^+etc.) (6)

From this general equation, the E.M.F. curve for any given spread of the armature winding
can be found. Several interesting cases are worked out in the paper referred to above.
Similarl}' from the general equation of the trapezium (see Fig. 323),

ar=^^(sin/8|sin0,+^sin3i9|sin3^,+ ...Y

the coefficients of the various terms in equation (4), page 306, are worked out. These are
especially interesting as they refer to the alternator with a C3'lindrioal field-magnet (see
page 377).

iSffect of the slots. The spacing ripple. If Z denotes the total number of
slots in the periphery, and Q the number per pole, then the slot-pitch in radians

will be : = ^ = ^ == y. The slot-pitch y then denotes the angle between successive

coil-sides. We now have to find the sum of the e.m.f.'s induced in the m coils
displaced from one another by the angle y (see Fig. 324). This depends upon the
value of the winding factors of the harmonics in the field-form.

Where the pole-pitch is exactly divisible by the slot-pitch, t.e. Q is integral, the actual coils
can be replaced by full-pitch coils (see Fig. 116). Then, from Eq. (3a), p. 306, for m full-
pitch coils in series:

2;;'B, = B«+B6+ . . .-t- Bm (see Fig. 324)

= Bi(sin^«-i-sin ^^-f ...-fsin ^«) (see Eq. (1), p. 22)

H- Ba(sin 3^a-i-8m 3^*+ ...-|-8in 3tfm),

-t-etc.

-fete.

• //. m-l \ . my . J^ ,m-l \ . ^wy

sinf ^a-f-— 2— TJsm-g- sinSI ^a-f-— 5-7 jsm3-2^

sin I sin3|

m. — 1 /I I a

Now, Oa + y = "q "* = 6 ssdisplftcement of midpoint of the m coil-sides ; hence,

sinm^ sin3m|

7:^e=2TcvllO-^ \B^ fsin^H-Ba sin 3^^- etc. \.

sin I sin3|

BinA

Since <m, the harmonics in the pressure Z^t of m coils in series will be less

sinA^

than m times the harmonics in the coil pressure.

310

DYNAMO-ELECTRIC MACHINERY

Inserting T=m. Te= total turns in series, the general expression for the pressure induced
in m coils at angle y apart becomes :

my . -m*)

Sin

jre=2TvLlO-^\ Bi

msin-

sin ^+-88

I y

msinS^

sin 39+..

=2Tt;L10--«(Bi/iSin9+BJ,sin3(?+...)=2!rt;il0-»Bi(/iSin9+|^/8sin39+...).

This is of the same form as Eq. 4 for distributed windings, but the winding factors are now :

Bin 3^

/i='

/.=

msm^

msbxZl

smA-^

. , wiir

''*^ Ty^ i V*

msinA^ msinA^ a

smce 7=7).

These are seen to be different from the winding factors for uniformly distributed
windings given on page 306, and we must now investigate the influence of this on
the shape of the pressure curve. In the present case Q is an integer, i.e. any odd

^ ^-^^/^-^/y - ^i

<?.

->J ...^ ...^^

■

FIG. 324.

or even number, so that 2Q — ^the number of slots in a double pole-pitch corre-
sponding with a complete period — ^will always be even. Further, in steady flux
curves, with the positive and negative parts identical, only odd harmonics are
present; hence 2Q±1, 2Q±3,..., 2Q±x and M.2Q±x will be possible values
of h for the harmonics B^, where x is any odd number and M any whole number.

For these particular harmonics, winding factors become :

/(2«-l)=/(2«+l) =

. 2Q±l T

= ±

m IT
Q 2

. 2Q±l IT ^ .It
ffism.-^-^—^- msm^ 5

= ±/i.

Similarly,

and
and

Q 2

/[2«-3) =/(a«+3) =±I»
f{iQ-x) =fm+x) = ± /x ,

f[M2q^x) =fiJi2q+z) = ± /«•

Thus when there is a whole number of slots per pole, the winding factor does not decrease as
the order of the harmonic h increases in the same way as with uniformly distributed winding,
but periodically rises to a maximum (numerically =/i) whenever h passes a multiple of 2Q« For
example, with Q = 6 or 2Q = 12, as in a three-phase winding with two slots per pole and phase
(m=g=2), we get /*=/(i^.2«-i)=/(jir.2«+i)= ±/i when ^ = 11, 13; 23, 25; 35, 37; 47, 49;
etc. (op. Table XVI.).

ALTERNATING-CURRENT GENERATORS

311

This means that if any of these hannonics are present in the B-cnrve, they will
reappear in the pressure curve ^e with the same percentage value as the
Aindamental, whilst the other harmonics are largely reduced by their winding
factors. Thus any of these harmonics will give rise to a ripple on the Amda-
mentaL

Since this e£Eect is due to the spacing of the armature coils, Dr. S. F. Smith
has given to it the term " spacing ripple" It is easy to see that this ripple in

(a) Coil presBure.

(b) Phase preasure (one leg of star).

(e) Tennlnal pressure (two legs in series).

Fig. S26.

the pressure wave will be mainly due to harmonics of the orders 2Q±l, since the
values of B^ become very small for M .2Q±1 when M>1. Also the harmonics
of the orders 2Q ± 3, with winding factors numerically equal to/j, will not be nearly

312

DYNAMO-ELECTRIC MACHINERY

80 important as the (2^ + l)th. Again, as the number of slots per pole increases,
B(2(2+i) usually decreases and the spacing ripple becomes less pronounced. For
example, in a three-phase winding with 6 slots per pole and phase, 2Q ± 1 = 35 and
37, and both Bgg and B^- are very small even with a rectangular flux distribution.

The oscillograms reproduced in Fig. 325 (a), (6) and (c) taken off a machine with
semi-enclosed slots, having three slots per pole per phase, clearly show the spacing

Fig. 326.

ripple. In this machine, the flux curve is fairly rectangular and smooth, as seen
in the curve marked " coil pressure," showing that there is practically no swinging
of the flux, so that the ripple is almost entirely due to the spacing of the coils.
The magnitude of the spacing ripple can easily be calculated when the harmonics

FIO. 327.

in the B-curve are known. Fig. 326 gives the three curves superimposed for a
machine having only two slots per pole per phase. Here the main indentations
on the terminal pressure are due to the " spacing ripple." Some slight " tooth
ripples " (see page 313) are visible in the flux and phase-pressure.

The winding factors for each phase of ordinary 3-phase open windings with
a whole number of slots per pole are given in Table XVI. To find the corresponding

ALTERNATING-CURRENT GENERATORS 313

winding factors for the terminal e.h.f., or the e.h.f. of a single-phase winding with
two-thirds of the slots wound, the values in the table mnst be multiplied by cos h 30'.

Table XVI, Wikding Factobs fob Phase e.u

3-PaAaE Windings in Slots.

«=

M__

B/>=J.

■966

960

■968

967

•957

•956

■965

-956

■707

667

■664

646

644

•642

•641

■640

■639

•636

■269

217

■206

200

197

•185

■194

■104

■193

191

-■2M -

177

- -168 -

149 -

145

-143

-141

-140

-136

- 707 -

333

- -270 -

247 1 -

236

-■229

-■225

-■222

-■220

- 212

/..

-«66 -

177

- -120 -

110 -

102

-■097

-■096

-■093

-■092

-087

- M6

217

■126

102

092

■086

■083

■081

■079

■073

- '707

667

■270

200

172

■168

■150

145

■141

-127

-259

960

■158

102

0»4

■076

■070

■066

■004

056

■269

980

-■206 -

110 -

084

-072

-068

-■062

-060

-■060

■707

667

-■664 -

247 -

172

-■143,

-■127

-118

-112

-091

966

217

-968 -

149 -

092

-■072 1

-■063

-■067

-054

-■041

■966 -

177

-■968

200

102

■076,

■063

■066

■062

■038

ft,

■707 -

333

-■664

646

236

■168

■127

■111

■101

■071

/»

■268 -

177

-■205

967

146

■086

■066

■056

■050

■033

The chief hfimionios in the spacing ripple, i.e. 2Q± 1, a

Bttvy type.

Where the number of slots per pole is fractional, the eSect of the spacing ripple
is veiy much reduced, as explained on page 306. Fig. 327 ia an oscillogram taken
from a machine having 6-75 slots per pole. Here the spacing ripple so apparent
in Fig, 326, where 0 = 6, has entirely disappeared, and we get for both the phase
pressure and the terminal pressure quite smooth curves as with uniformly dis-
tributed windings.

Pnlaations due to the teeth. The tooth ripple. When a number of teeth
like those depicted in Fig, 328 are rotating under a pole they produce pulsations
of the flux of two kinds : (1) a pulsation in the total flux per pole, due to a change
in the reluctance of the air-gap as the armature changes from position (a) to position
(6) ; and (2) a swinging to and fro of the flux along the periphery as a tooth under
the horn of the pole is replaced by a slot. Both of these effects can be very much

314

DYNAMO-ELECTRIC MACHINERY

diminiBhed by isuitably bevelling the pole so that the reluctance under the polar
horn IB almost constant for any position of the armature. The skewing of the
slots or of the polar horn by a full slot pitch (see Fig. 533) is also a cure.

We do not in practice find very much pulsation in the amount of the total
flux per pole, because such a pulsation would be opposed by eddy-currents in
the solid parts of the magnetic circuit, and by alternating currents induced in
the exciting circuit. The swinging of the flux, however, may give rise to very

/jitfs^

Fig. 829. — Calculated tooth ripple due to swing of sinuBoidal flux. Q=9 slots per pole.

noticeable ripples in the wave-form. Smith and Boulding employ the name of
" tooth ripples " for these, to distinguish them from the " spacing ripples " described
on page 310.

When a tooth approaches the horn at the right side of a pole, it reduces the
reluctance of the air-gap between itself and the pole, and the fringing flux extending
to it rapidly increases. At the left side of the pole a slot may be taking the place
of a tooth, so that the fringing flux on that side is rapidly diminishing, and the
disposition of the flux becomes unsymmetrical with regard to the centre line of
the pole. After the movement of the armature through one-half a tooth-pitch,,
the flux is again unsymmetrical ; but now the heavy fringing is on the left side
and the lighter fringing on the right. Such a swinging to and fro of the flux gives
rise to b.m.f.'s in all the conductors under the pole, which are superimposed upon
the E.M.F.'s generated by the uniform movement of the conductors. The sum of
the eflects in all the conductors in a phase-band is greatest when B is greatest, that is

ALTERNATING-CURRENT GENERATORS

315

when the phase-band is opposite a pole, and is least when the phase-band is
between the two poles. Thus the ripples due to swinging of the flux are greatest

(a) Coil pieBnire.

(&) Fhaae prenure.

(e) Tenninal pressiire.

Fio. 330. — Oscillograms taken on 3-phase machine having 6 open slots per pole, and showing

the " tooth ripple ** in a marked degree.

on the crest of the wave, and sink to a minimum as the main wave passes through
zero, as will be imderstood by reference to Fig. 329. The exact shape of the ripples
is, of course, very complex ; but as a first approximation we may take them as

316 DYNAMO-ELECTRIC MACHINERY

sinusoidal. On this assumption, it can be shown that where the flux from the
pole is sinusoidal, the ejBEect of the tooth ripple can be expressed by the addition
of a term to the ordinary expression for the e.m.f. in a coil. Thus the e.m.f. in
a coil of Tc turns becomes :

6 = 29rnrc<^ 10-8 (sin ^1 -ly 2Q sin ^1 cos 2g6'i)

where f] is the amplitude of the tooth ripple, and Q is the number of slots per pole.

Now, -2sin 6^1 cos2Qei=sin(2Q- 1)^1 -sin (2g + l)^i,

so that the tooth ripple can be analysed into two odd harmonics, the (2Q-l)th
and the (2Q + l)th, each having an amplitude equal to half the maidmum amplitude
of the ripple as seen on the crest of the wave. In Fig. 329 are shown the tooth
ripples occurring in a machine having 9 slots per pole : 2Q = 18. Along the base-line
are plotted the seventeenth and nineteenth harmonics, which when combined
give the characteristic shape of the tooth ripple.*

Fig. 330 gives the coil pressure, the phase pressure and th,e terminal pressure
of a 3-phase machine having 6 open slots per pole. The fact that the flux was
swinging is shown by the ripples on curve (a). These tooth ripples appear un-
diminished in phase pressure (6), and the terminal pressure (c). The dissymmetry,
of the curves is probably due to the hysteresis of the iron of the armature.

Where the phase-band is made up of m coils connected in series, the amplitude
of the ripple in the phase pressure 2*6 depends upon the value of the winding
factor of the (2Q - l)th and {2Q + l)th harmonics. Now, it was shown on page 310
that when the fleld system is normal and Q is integral :

/(2<3 - 1) =/(2Q+l) = ±fu

so that in this case the tooth ripple occurs in the phase pressure with the same
percentage value as in the coil pressures.

Where the number of slots per pole is fractional, the effect is reduced just in
the same way as with the spacing ripple. It is as if the number of slots per pole
were increased (see page 305).

CALCULATION OF A 750 K.V.A. ENGINE-DRIVEN ALTERNATING-CURRENT
GENERATOR, TO RUN AT A SPEED OF 375 REVS. PER MINUTE.

2100 VOLTS ; 3 phases ; 50 cycles.

In going through the calculation given below, it may be convenient to refer
to Fig. 331, which shows the generator in question drawn to scale. Fig. 332
gives a sectional elevation, and Fig. 333 gives details of the poles and field-coils.

We will suppose that we have obtained an order for a 750 k.v.a. generator
to comply with Specification No. 1. The size of the frame upon which such
a machine would be built would depend upon the particular sizes of frames which
the manufacturer might already have ; but if we were to start de novo, the con-
siderations which would settle the diameter and length are those given on page 299.

* For discussion on question whether the tooth ripple is symmetrical, see remarks by C. C
Hawkins and Dr. G. W. O. Howe,«/aum. I.E.E., vol. 53, pp. 241, 243 ; also EUctrician, vol. lxxiii.»
pp. 3, 367. 417, 466, 497, 637.

ALTERNATING-CURRENT GENERATORS 317

The higher the peripheral speed, the better the specific use we make of our
copper and iron ; but if we choose too high a peripheral speed by making the
diameter great, we find that the axial length of the generator becomes short as
compared with the pole pitch, and the cost of construction comes out higher than
one would at first suppose from the mere statement of the weight of active material.
Moreover, if we make the peripheral speed much higher than 6000 feet per minute,
it will be found necessary to adopt a special construction of field spider in order
to provide against the great centrifugal forces. A peripheral speed of 6000 feet
per minute for a 50-cycle generator gives a ratio of pole pitch to axial length which
is very economical, and no expensive construction of field-frame need be resorted to.

We will therefore decide on a peripheral speed of 6000 feet per minute, or say
30 metres per second ; and we will take the internal bore of the stator as 155 cms.
In the calculation sheet given on page 321 the dimensions are given both in inches
and centimetres.

This calculation sheet is designed so that the same general form can be used
either for an A.C. generator, a c.C. generator, an asynchronous motor, a synchronous
motor, or a rotary converter. As explained above on page 8, the same general
method will be used when calculating all these machines, and it is an advantage
to be able to compare the figures of one type of machine with those of another
type on the same form. The first three lines of the form in question deal with
the performance of the machine which it is intended to design ; after the date
comes a statement of the type of machine, whether a turbo- or engine-type, belted
or geared, open or enclosed, etc. The number of poles is of course an important
matter, which naturally takes a conspicuous place. Then comes the electrical
specification number, in one corner for easy reference. The second line is self-
explanatory. In the third line it is necessary to insert the amperes per conductor,
because sometimes there are several paths in parallel through a machine. For a
■ C.C. machine or rotary converter the amperes per brush arm should be stated. The
temperature rise and regulation and over-load capacity are all matters relating to
the guara£iteed performance. The next line deals mainly with records, such as
the name of the customer, the order number, the quotation number where a quota-
tion has b^en made previous to the order, and the performance specification num-
ber. The fourth and fifth lines deal with important data belonging to the size
of frame employed. The frame number, which is sometimes specified by giving
.. the diameter of the bore and the length of iron. The amount of air required for
cooling, if the machine is of the turbo-type. The circumference of the active
surface ; the gap area or area of active surface ; the ^^B, for which see page 6.
Above this it is convenient to write the greatest possible AgB that can be put on
the frame in question. The /oZa, for which see page 8, and the greatest possible

Zo^a* The — — ^^ — - gives the ampere wires per centimetre of periphery, a very

important quantity in judging the rating of the frame. Then comes the output
coefficient. The next line gives the £«, for which see page 23 ; the voltage formula ;
the ampere-turns per pole on the armature ; and lastly, the total maximum ampere-
turns on all the poles. The left-hand side of the form then deals with the arma-
ture, which may be either revolving or stationary ; and the right-hand side deals

or, !100 rolU, £0 erdea,

I :

FIO. 382.
875 R.P.if., designed to meet Specification No. 1, page 269.

320 DYNAMO-ELECTRIC MACHINERY

with the field-magnet, which may be either stationary or revolving, the word that
does not apply being struck out in each case. The general method of using the
form will be best understood from the example given below. Two columns are
provided for the insertion of figures on each line. The purpose of these two columns
is not, as might be supposed from an inspection of the form on page 321, for the
insertion of both centimetre and inch units ; for each engineer will as a rule con-
fine himself to the system of units which he prefers. The second column is to
enable the figures ioz an alternative design to be put alongside those of the principal
design, in order that comparison may be conveniently made. We have used the
second column on page 321 for the insertion of the dimensions expressed in inches,
for the convenience of such readers as are more familiar with those units. The
rough diagrams of the slots, teeth and poles are on the original' form drawn so
that by means of a few simple lines it is easy to represent either open slots or semi-
closed slots, and various shapes of pole.

We will proceed, then, to fill in our calculation sheet for the 750 K.v.A. engine-
type generator to give a three-phase current (power factor 0-8) at any voltage
from 2000 to 2100. The amperes per terminal will be 206 ; the cycles per second,
50; the R.P.M., 375. The amperes per conductor in this case would be 206. Accord-
ing to the specification on page 270, the temperature rise by resistance after full-
load run will be 55° C. ; the regulation on unity power factor 8 per cent. ; and
the overload, 25 per cent, for two hours. If ^^e adopt an inside diameter of stator
punchings of 155 cms., we arrive at a circumference of 486 cms. The final fixing of
the exact length of iron cannot be done until the design has proceeded somewhat
further, but a preliminary length can be worked out from what we know to be a
suitable D^l constant. For a high-speed engine-driven generator having the
performance specification on page 270, a suitable DH constant would be 4 x 10* cm^. ;
this would give us about 32 cms. length of iron ; and if one of our standard frames
happened to be 12J ins. long, or 31 -8 cms., we would make an attempt to get the
machine on that frame. Multipljdng the circumference by 31-8, we get Ag, the
gap-area or area of active face equal to 15,400 sq. cms. If we could have a flux
density in the gap of 10,000, this would give us a possible AgB of about 1 -5 x 10^.
Applying the voltage formula ((1), page 24), we would arrive at a total number of
conductors of about 592. But it is convenient to have the number of conductors
a multiple of 48 or 16 times 3, so a more suitable number is 576. We can then
have 144 slots with 4 conductors per slot. 144 slots gives us 9 slots per pole, or
3 slots per phase per pole. We ought to say something here about the considera-
tions which settle the number of slots per phase per pole. The cheapest arrange-
ment of conductors is of course one which employs a small number of slots ;
because it is necessary to insulate each coil for full voltage to earth, and as we
increase the number of coils, we increase the space taken up by insidation as well
as the cost of the insulation. If we were to employ only 1 slot per phase per pole
in the machine under consideration, we should have about 2500 amperes per slot.
The amount of heat generated per coil would be three times as great as in the
arrangement proposed, and as the cooling surface of the coil would not be increased
in proportion, the heating would be excessive; unless, indeed, a greater cross-
section of copper were used. Moreover, the wave-form of the generator would

ALTERNATTNG-CUBRENT GENERATORS

D^,ift*»»Wi,«. rjfMT,.... .^/*,.i^.cwu

■***• Asp* p. aat-^Pff. *■!■ n V ■■■ i^.mf,.rmv »*■» tff-'tH--

CrdM..^^. -J RJ>.II.iK^- I KMh «w.. ~

c«=.>„ ^5r«w .^fi«.5«. : 0^ ^.S4iL. : Q^ naMSl..^ P-t ^-A^aM--

322 DYNAMO-ELECTRIC MACHINERY

be of very irregular shape. Two slots per phase per pole would be a possible
arrangement ; but this would give as much as 1650 amperes per slot, a rather high
figure for a small machine of this type. Three slots per phase per pole give us
better cooling conditions, and at the same time a very smooth wave-form. If we
were to try to put in 4 slots per phase per pole, we should find that the amount
of room taken up by the insulation, in comparison with the room taken up by
copper, would be excessive.

We shall therefore decide to have 576 conductors, there being 144 slots with
4 conductors per slot. Filling 576 in the voltage formula, we obtain :

2100 = 0 4 X 6 -25 X 576 x AgB,

AgB = 1 -46 volt-lines.

On 50-cycle generators of this type it is well to have a ventilating duct for
every 5 cms. of iron (see page 254). This will give us, say, 5 ventilating ducts,
each 0-635 cm. wide. The net length is then obtained by multiplying 31-8 -3-2
by 0-89 : thus we get 25-4. To see whether this is sufficient, we must fix upon
the size of slot, and this depends upon the size of conductor to be employed. The
final fixing of the size of conductor will depend upon the cooling conditions of
the armature coil ; but as a preliminary figure we may, for an armature of this
kind, assume 380 amperes per sq. cm. This suggests a conductor of a size
0-75 X 0*75 cm., having an area of 0-53 sq. cm., allowing for the rounded comers.
A suitable thickness of insulation between copper and iron, consisting of mica and
manilla paper, for a 2100- volt generator, is 0*2 cm. (see page 202). Adding the
double thickness, 0-4 cm., to 0-75 cm. of copper, we arrive at 1-15 for .the net
thickness of copper and insulation. We should add to this an additional allow-
ance of 0-12 cm. for the staggering in the building-up of the punchings and for
air spaces. This gives a total width of slot of 1 -27. In calculating the depth of
slot required, we must not forget that it is advisable to place built-up mica between
each conductor : so that 4 conductors and their insulations would take up
3-4 + 0-44 cms. ; and allowing an additional 0-35 cm. for a fibre wedge, we arrive
at 4-2 cms. for the length of slot.

We can now proceed to find the maximum flux-density in the iron teeth. This
we do by dividing the total section of all the teeth into the total AgB* As the
sides of the slots are parallel, the sides of the teeth will not be perfectly
parallel: so that the density of the flux will not be uniform all along the
teeth. In cases where the change in the flux-density is very great, it is desirable
to adopt a special method for considering it (see page 73) ; but in cases of
this kind, where the diameter of the armature is great as compared with the
depth of the slots, it is sufficient to take account of the flux-density at a point

•According to the older method of calculating a.c. generators, in which the flux per twle
is taken as the quantity from which all flux-densities are calculated, it would bo usual to
estimate the number of teeth per pole and divide the area of tlie cross-section of these teeth
into the total flux of the pole. The number of teeth per pole is a quantity which we cannot
l)e very certain of where the polo is bevelled or where tne flux-density tails off towards a
neutral line. It will be rememl)ered that the quantity AgB is arrived at by multiplying the
maximum B by the whole area of the gap. If, therefore, we divide the whole area of the
teeth into the quantity ^^B, wo arrive at the maximum density of the teeth at no load, or
at any other load for which we arc given the maximum flux-density in the gap.

ALTERNATING-CURRENT GENERATORS 323

one-third of a tooth length from the narrowest part of the tooth. This can be
obtained with sufficient accuracy by the following method : for teeth eicternal to
the air-gap add to the diameter 0*66 of the length of the teeth, multiply the sum
by TT, and thus obtain the circumference of the mean circle drawn around the
machine, passing through points one-third of a tooth length from the narrowest
part. The circumference in this case is 500 cms. Subtract from this the total
width of all the slots, 144x1-27 = 183. This gives us a total width of the teeth
of 317. Multiplying by the net length, 25-4, we arrive at 8000 sq. cms. for the
section of all the teeth. Dividing this into 1 -46 x 10®, we arrive at 18,300 C.G.s.
lines per sq. cm. As this is not an excessive flux-density for the teeth of a 50-cycle
generator, we may flx on the gross length of 31 -8 cms. as suitable. Referring to
the iron loss curve (Fig. 29), we find that the loss per cu. cm. of iron is 0*15 watt.
Now the volume of all the teeth is 8000x4-35=35,000 cu. cms., giving a total
loss in all the teeth of 5250 watts. We will now consider the depth of core below
the slots : this will in general depend somewhat upon the standard size of frame
and the depth of the slots. It is sufficient in a 50-cycle machine to provide such
a depth that the flux-density does not exceed 12,000 C.G.s. lines per sq. cm. In
this case we have taken the outside diameter of the punchings at 184 cms. ; this
gives a depth below the slots of 1015 cms., a cross-section of 258 sq. cms., and a
volume of 142,000 cu. cms. As the loss per cu. cm. is 0-06 watt, the total loss
behind the slots is 8500 watts. We will now return to the armature conductors.
We may take the length for the slots at approximately 32, and a length of end-
connectors of 62, giving a sum of 94 cms. ; so that 576 conductors would give a
total length in all phases of 540 metres. The calculation of the total weight of
armature copper is most easily carried out without reference to any wire table
by remembering that 1000 metres of copper wire having a section of 1 sq. cm.
weigh 875 kgs. If, therefore, we multiply 875 by the section of the conductor,
in this case 0-53 sq. cm., we obtain the weight 464 kgs. per 1000 metres ; so that
540 metres weigh 250 kgs. To obtain the resistance of any wire per 1000 metres at
20° C, we have the rule : divide 0-174 by the cross-section in sq. cms. 0 174 — 0-53
=0-328 ohm per 1000 metres ; so that the total resistance of all phases is 0-177
ohm. The total PR loss in the armature at full load will be 0-177 x 1 -2 x 206 x 206
= 9000 watts. For the calculation of the cooling of the copper in the slots it is
generally convenient to take the total loss in 1 metre length of coil : this is equal
to 0 000328 X 1 -2 X 206 X 206 X 4 = 67 watts per metre. The surface presented by
the insulation works out to 1050 sq. cms. per metre length, giving 0-064 watt per
sq. cm.

In order to find out whether the insulation can conduct heat at the rate of
0-064 watt per sq. cm. with a reasonable difference of temperature between the
copper and the iron, one should work out the heat conductivity of the insulating
tube just as it is done in the example given on page 222. Taking the conductivity
of the pressed paper and mica at 0-0012 watt per sq. cm. per degree, and the
thickness of the insulation at 0-25 cm., we will have for 15° C. difference of tem-
perature between copper and iron

0 0012 X 15
—0^25 ^^^-

324 DYNAMO-ELECTRIC MACHINERY

We see, therefore, that we have quite sufficient cooling surface on the insulating
tube to get rid of the heat generated within the coils. The cooling of the ends of
the coils depends upon the shape of the coils, the amount of space aJlowed between
each coil, and the velocity and temperature of the air circulating around them.
Usually the circumstances are too complex to permit of any calculation, but ex-
perience shows that if the individual coils are kept separate, as shown in Figs.
331 and 114, so that the air can blow in between them, the cooling conditions
for the end windings are at least as good as for the parts Jying in the slots.

Specification No. 1 requires that the armature coils shall be able to withstand
a short circuit. In Chapter VI. we considered the forces which come into play
when a machine is short circuited at full voltage. It is easy to show from the
considerations there taken up that the forces on the coils of this machine are not
very great ; and as the average throw is only 31 cms., the coils themselves, if
bound together as shown in Fig. 331, are sufficiently stiff without attachment to
any framework. In the case of 25-cycle machines, however, where the throw of
the coils is greater, and where the ratio of the leakage flux to the flux per
pole is only half of what it is in this case, the danger to the coils is considerably
greater ; and it is well in big generators of this kind to brace the coils by means
of special clamps, as shown in Fig. 113&.

Cooling of the stater. It now remains to add up all the losses occurring in the
stator, the heat from which must be dissipated from its surfaces. It is usual to
assume that those parts of the stator coils which project into the air will be cooled
by the draught of air blown upon them, so that only that part of the PR armature
losses which is produced in those parts of the coiJs lying in the slots, " the buried
copper," need be taken as adding their heat to the total heat dissipated by the
stator surfaces. The total " buried copper " losses are readily calculated by
multiplying the watts per metre by the total length of all the slots. This gives
us about 3000 watt^. Adding together the loss behind the slots and the loss in
the teeth, we get 13,850, which with 3000 gives us a total loss of 16,850 watts. It
is now necessary to see whether the cooling conditions of the stator are sufficiently
good to get rid of the loss with a temperature rise not exceeding 45° C.

There is a rough-and-ready method which is sometimes used to get a rough idea
of the total amount of heat which can be dissipated by the stator. According to
this method, one adds the total external surfaces of the stator to the surfaces of the
ventilating ducts, only one side of each duct being counted as effective. One
then allows so many watts per sq. cm., the allowance being based upon the observed
temperatures of similar machines running at about the same speed. For generators
and motors of the ordinary type having a peripheral speed of 6000 feet per minute,
or 30 metres per second, one may usually allow 1 watt per sq. in., or 0-155 watt
per sq. cm. This method, though somewhat crude, gives sufficiently good results
if we have means from time to time of correcting our coefficient. Appljdng it in
our present case, we get as the total cooling surface 87,000 cu. cms., so that the
total watt« dissipated for 45° C. rise lies between 16,000 and 18,000.

A more accurate method is to apply the rules given in Chapter X. in esti-
mating the watts dissipated from the inside cylindrical surface, or gap-area, the
vent-area and the outside area respectively.

ALTERNATING-CURRENT GENERATORS 325

Watts dissipated from gap-area. The peripheral speed of the rotor is 30 metres
per second, and from formula (1) (page 230) we have

ntfo ri 333 X watts per sq. cm.

^ ^- — iT3 •

watts per sq. cm. =0-42 ;

so that the gap-area can dissipate 042 watt per sq. cm. We have taken the
difference in temperature between the iron and the air in the air-gap at 35*^ C.
This allows 10° margin for the heating up of the air in the gap. On a total area of
14,000 sq. cms. we get rid of 5900 watts.

Watts dissipated from vent-area. To arrive at hv, one should know the velocity
of air in the ventilating ducts. In enclosed machines with definite air channels
this velocity is fairly well known. But in open machines it depends upon so many
factors that it is difficult to estimate it even approximately. Where, however, we
employ well-shaped vent spacers, and where we have plenty of room between the
coils of the rotating field, we may take the mean velocity of air in the vents at
one-tenth the peripheral speed of the rotor. In this case we may take it at 3
metres per second. This gives us A,, = 0 0042 watt per sq. cm. per degree C. rise.
It then becomes necessary to make a rough estimate of the difference between the
mean temperature of the air in the ducts and the temperature of the iron. If we
take the mean temperature rise of the air entering the ducts at 10° C, and of the air
expelled from the ducts at 30° C, and taking the surface rise of the iron at about
40°, we have a mean temperature difference pf 20°. Multiplying this into 0-0042,
we arrive at 0 084 as the watts per sq. cm. dissipated from the ventilating ducts.
Multiplying by the area of the ducts counting both sides, 75,000 sq. cms., we arrive
at 6300 watts dissipated from the ducts.

Watts dissipated from the external surflftce. A great deal of heat is conducted
from the punchings to the cast-iron frame, whence it passes by convection and
radiation to surrounding objects. It is impossible to make an accurate estimate
of the amount of heat lost in this manner. A simple plan, which gives sufficiently
correct results in practice, is to take the total external surface which is made up
of the two end-plates and the external cylindrical surface, and to multiply by
the coefficient 0-15 watt per sq. cm., which is equivalent to about 1 watt per sq. in.
The total external surface in this case amounts to 33,000 sq. cms. dissipating
4950 watts.

Estimating the cooling in this way, we arrive at a total figure of 17,150 watts
dissipated by the stator iron sitrfaces for a temperature rise of 45°.

Design of the field-magnet. We now come to the design of the field-magnet.
The considerations which govern the number of ampere-turns required on it in
order to get the desired regulation have been dealt with on page 278.* We have
seen on page 276 that there are reasons for making the pole wide at the base, while
there are other reasons for making it narrow immediately below the polar horns.
In order to avoid a taper pole, we may make the pole body in two parts (see

♦ And see " Effect of Leading and Lagging Currents on Regulation of Alternators," B. N.
Westcott, Elec. World, 89. p. 46, 1912 ; " Regulation of Definite-pole Alternators," Mortensen,
Amer. In$t, E.E., Proc. 32, 291, 1913; "Experimental Determiaation of the Regulation of
Alternators," Field, Amer, Inst, E.E., Proc. 32, 599, 1913.

326 DYNAMO-ELECTRIC MACHINERY

Fig. 334), each of rectangular shape, which are held together by the bolts which hold
the pole to the main ring of the field-magnet. The making of the poles in two
parts and the winding of the coils in two parts increases the cost of labour by a
very small percentage, and enables the output of the frame to be increased by
about 8 per cent.

It will be seen from Fig. 334 that this construction allows a cross-section of
540 sq. cms. at the root of the pole, while imnaediately under the horns of the
pole the section is only 379 sq. cms. For a length of 8-3 cms. we are able to get
sufficient saturation to improve the regulation of the machine without nmning
any danger of excessive saturation at the root of the pole when the generator is
working on heavy load of low-power factor.

The dimensions of the air-gap. As we have seen on page 62, the fixing of the
length of air-gap depends upon a number of considerations. In the first place the
air-gap must not be so small that the unbalanced magnetic pull for a small acci-
dental displacement of the rotor is excessive. In practice it will be found that
it is only on machines of very large diameter with a great number of poles that this
consideration has much weight in controlling the width of the air-gap (see page
347). In machines of smaller diameter the air-gap has to be made fairly wide
in order to get sufficient ampere-turns on the pole to obtain the desired regulation,
and it is then found that the unbalanced magnetic pull is not excessively great.
On page 278 we have given the relations which must exist between the ampere-
turns on the field-magnet and the ampere-turns on the armature, in order that an
alternating-current generator may possess certain regulating qualities. In the
case of the 600 K.w. machine under consideration, the effective armature ampere-
turns per pole are 3800. To obtain not more than 8 per cent, rise in voltage when
full non-inductive load is thrown off, it will be necessary to have about 5600 ampere-
turns per pole at no load, even with considerable saturation of the iron. This
we know from the construction given in Fig. 310, though the final adjustment of
the ampere-turns per pole can only be arrived at by a process of trial and error.
It is thus found that an air-gap of 0-2 in. or 0-51 cm. will be sufficient to give the
desired regulation.

The magnetic flnx per pole is found from the AgB by the formula,

TT — f-^^— =flux per pole (1)

number of poles r r \ /

Tr, ¥\\^ naao 1 '^^ X 10® X 0-655 ^ ^^ ,^

In this case _ = 5 -96 x 10®.

Oalcnlation of leakage * between poles. Before we can estimate the number of
ampere-turns absorbed in driving the flux along the body of the pole at no load
and at full load, it is necessary to make a calculation of the leakage flux. This is
best done by means of a graphic construction such as that given in Fig. 333.
The procedure is as follows : First lay out a vertical line to represent to scale an
imaginary line drawn along the neutral plane between the poles in a radial direc-
tion. This line in our present case will be 20 cms. long. Then draw a diagram

♦ The reader should also consult a paper by Dr. Pohl, Jour. Inst, Eke. Engrs., voL 52, p. 170,

ALTERNATING-CURRENT GENERATORS

327

which gives the distance from the iron of the pole to the neutral plane, as shown
by the thin dotted line in Fig. 333. Then set ofiE a curve, the abscissa of which
gives the magnetomotive force exerted by the field-coil between the iron of the
pole and the neutral line. At the root of the pole this magnetomotive force will

t4eoo M.M.F.

o noo 2000 3000 FLux.- density CMS. units

Fig. 333.— Construction for finding the leakage per pole.

be zero, and as we pass radially outwards along the neutral plane, the magneto-
motive force will increase, the rate of increase depending upon the number of
ampere-turns per cm. length of pole. In those cases where the field-coil is rect-
angular in section, so that the ampere-turns per cm. are constant, the magnetomotive
force curve is a straight line. We obtain the extreme comer of the curve by multi-
' plying the ampere-turns per pole by 1-257, and plotting a point which has the

328 DYNAMO-ELECTRIC MACHINERY

value thus obtained for its abscissa and a vertical height equal to the height of the
coil for its ordinate. In our case the winding on the pole consists of two rect-
angular coils having different numbers of turns per cm., so that the magnetomotive
force at no load will be represented by the dotted curve marked " mjic.f. no load,"
which has two straight sections of different slopes. The vertical part of the curve
shows that the magnetomotive force is constant for all points on the neutral plane
beyond the limits of the coil. In those cases where the number of ampere-turns
absorbed by the pole itself is small, it is sufficient to use the curve of magnetomotive
force yielded by the coil to obtain the flux-density between the poles. But in a
case like the present, in which there is a deliberate intention to absorb a considerable
fraction of the ampere-turns on the pole itself, one ought to deduct from the value
given by the coil magnetomotive-force curve a certain amount for the ampere-tums
lost at each point along the neutral line. For instance, in this case 750 ampere-
tums (M=940) are lost in the 8 cms. of pole body, so it is necessary to draw a new
M.M.F. curve. This is shown by the thin full line in Fig. 333. In order to get
the flux-density between the poles, it only remains to divide the effective magneto-
motive force at each point by the distance from the iron to the neutral line at each
point. This gives the curve shown by the thick line. This curve can be plotted
fairly definitely in the upper reaches between the tips of the pole, and also in the
lower reaches near the root of the pole. The middle part can be filled in by an
easy-flowing curve, which we can draw by exercising some judgment upon the way
that the flux would spread along the irregular path which exists between the tips
of the pole and the centre of the pole. Having obtained this curve for the approxi-
mate flux-density at each point along the neutral line, we can find the mean value
either by the use of a planimeter or by any other method of finding the mean height
of the ordinate of a curve.

A rough-and-ready rule, which works very well in practice, for obtaining the
effective area of the path between the poles is as follows : Add one-third of
the pole pitch pp to the effective axial length of the pole ley and multiply this by
the total height of the pole hp . For instance, in Fig. 334 we may take

^Pp = 10 cms.,

^^ = 30 cms.,

hp = 20 cms.

20(10 + 30) =800 sq. cms. at each side of the pole.

If now the mean flux-density between the poles is 525 C.G.s. lines per sq. cm.
at no load, the no-load leakage may be taken at

800 X 2 X 525 = 0-84 X 10^ C.G.s. lines.

If the mean flux-density between the poles at full inductive load is 990 C.G.S.,
the leakage per pole will then be 1-58 x 10 lines per pole.

Adding the leakage to the no-load working flux, 5-96 x 10^, we get 6-8 x W for
the total flux per pole at no load, or adding 1 -58 x 10^ we get 7-54 x 10^ for the flux
per pole at full load.

With the type of pole illustrated in Fig. 334, we are able to carry the narrow
shank immediately under the pole piece to a higher flux-density than we would
risk at the root of a pole of the ordinary shape. We may safely employ a flux-

«-

00-

r»-

«-

ift-

t*-

10-

O —

N-

— -

10-

o—

00-
0>-

JBD
1

e
1

«
CO

330 DYNAMO-ELECTRIC MACHINERY

deDsitv of 18,000 at no load. This would give ua an area of 379 sq. cms. The
flux-density at full load would then be as high as 19,800. This high saturation
improves the regulating qualities of the machine.

Magnetization cnnre. It is usually sufficient to calculate the number of ampere-
turns per pole required for three different voltages, the normal no-load voltage,
another voltage some 5 per cent, higher and another voltage 10 or 15 per cent,
lower. It will be found in general that from the three points in the magnetization
curve thus obtained the curve can be drawn with sufficient accuracy.

Ampere-turns on the core. The number of ampere-turns on the core in 50-cycle
generators is usually so small that it can be neglected in view of the errors which
are sure to arise in the estimation of larger quantities. In the form given on
page 321, however, the ampere-turns on the core are calculated merely to illustrate
this fact.

Ampere-turns on the stater teeth. The section of all the teeth as given above
is 8000 sq. cm. The length may be taken as 4-3 cms. The flux-density at 2100
volts is 18,300, and by placing the number 2 1 of the movable scale of our slide-
rule opposite 18-3 on the fixed scale, we obtain B = 19,200 at 2200 volts, and
B = 15,700 at 1800 volts. Referring now to our magnetization curve, suppose we
obtain the figures 24, 97 and 160 for the ampere-turns per cm. in soft sheet steel
at the three given densities. Multiplying by 4-3 cms. length of tooth, we arrive
at 103, 415 and 690 ampere-turns on the teeth.

Ampere-tnins on the gap. We have seen that the length of the air-gap has
been fixed at 0-51, so as to absorb the required number of ampere-turns. The
flux-density in the gap at 2100 volts is obtained by dividing ^^B, 1 -46 x 10®, by
the gap-area 15,400 sq. cms. This gives us B = 9500. By proportionality we get
the figures 8150 and 9950 at 1800 volts and 2200 volts respectively.

Working out the gap coefficient Kg, according to the rule given on page 65,
we arrive at jBl^ = 11. The ampere-turns on the air-gap are therefore 3640, 4250
and 4450 at the three voltages.

Ampere-turns on the pole body. At no load the flux-densities in the pole
body will be 15,900, 18,500 and 19,400 for the voltages 1800, 2100 and 2200 respec-
tively. The length of the pole body is taken as only 8-3 cms.; that is to say, the
length of the shank directly under the pole face. The ampere-turns on this part
amount to 170, 200 and 210 respectively. The ampere-turns on the remainder of
the magnetic circuit may be neglected. Thus the total ampere-turns per pole
at no load amount to 4243, 5760 and 6760 respectively.

In plotting a magnetization curve for any particular frame, it is better to take
as ordinates the flux-density in the gap than the voltage generated in the armature.
It will be found in practice that many machines built on the same frame will be
carried to approximately the same flux-density in the gap, whereas the voltage
generat-ed in the armature will vary widely, depending, of course, upon the number
of conductors. One magnetization curve plotted with flux-density as ordinates will
do for any number of machines built on the same frame. We have accordingly
adopted this plan in Fig. 301.

In plotting the magnetization curve, it is well to draw first the air-gap line.
We take the point which represents the flux-density of 9500 and the ampere-turns

ALTERNATING-CURRENT GENERATORS

331

of 4250, and draw a straight line passing through it and the origin.- We then plot
the points given by the no-load ampere-turns at 1800 volts, 2100 volts and 2200
volts. Thus we obtain the no-load saturation curve shown in Fig. 301. In order
to obtain the dotted curve marked " Increase due to leakage on load," we must
obtain at various voltages the extra ampere-turns required for the pole shank when
the leakage is increased by the extra ampere-tums required on the pole at full
load. The method of arriving at these extra ampere-tums will at once be under-
stood from the table given below.

1800 volts.

A.T.

2100 volts.

2200 volts.

A.T. p. cm.

B A.T. p. em.

A.T.

A.T. p. cm.

A.T.

Pole body (full load),
Pole body (no load).

17000

50 ,

410
210

19800 205

1700
750

20800

300

2500
1250

Diflference,

200

950

1250

It should be noted that when dealing with a very highly saturated pole, as
in this case, some flux will be carried by the space occupied by the coiJ, and to
find values for the actual flux-density in the pole one must roughly work out the
value of Ks (see page 71), and make use of Fig. 46. In this case j^« = l-5, so that
at 2200 volts an apparent flux-density of 21,000 in the pole body means an actual
flux-density of 20,800.

In Fig. 301 we set off the difference 950, as shown by the line NN\ and the
differences for the other parts of the curve, and so obtain the increase-due-to-
leakage curve.

Having obtained these two magnetization curves, it is possible to calculate
with fair accuracy the ampere-tums per pole required at full load by the method
described on page 293. This method simplifies down to the construction given in
Fig. 310.

The calculation of the field winding. The shortest method of finding the
size of wire and the number of turns required is based upon a knowledge
of the number of sq. ins. or sq. cms. of coil surface required to dissipate the
heat from 1 watt lost. When a designer is frequently dealing with coils of
about the same size and shape, he finds from experience the amount of surface
per watt to allow. For coils of about the size and shape depicted in Fig. 331,
running at a speed of 5600 feet per minute, an allowance of 1 *35 sq. ins. per watt
will be ample, if in taking the surface we count the surface of the outside, the inside
and the ends of the coils. More exact rules for determining the cooling constants
of revolving field-coils are given in page 233, and a further example is worked
out on page 352 ; but in this case we will assume that the constant, 1 *35 sq. ins.
per watt over the total surface, is known.

The total surface of all the coils works out at 11,400 sq. ins., or 73,500 sq. cms.,
so that when running at full speed we can dissipate 8500 watts.

Consider next the exciting voltage, in this case 125 volts, and allow some margin,
so that even at full load we will still have some 20 per cent, or so on the rheostat.
In this case we have IR in the winding 104 volts. Divide this into the 8500 watts.

332 DYNAMO-ELECTRIC MACHINERY

We will require .about 81 -5 amperes exciting current, so that we will require about
129 turns per coil to get 10,500 ampere-tums. These turns may be divided between
the two parts of the coil as follows : 75 turns in the upper and 54 turns in the
lower. The lengths of mean turn come out 1-07 metres and 1-12 metres respec-
tively, giving total lengths of 1280 and 970 metres respectively. The required
resistance (hot) is obtained by dividing 104 by 81-5. It should come out about
1-28 ohms. From this, knowing the length of wire, we can determine the size
of wire. In the machine under consideration, owing to the cooling conditions
on the respective parts, we ought to make the upper part of the coil in Fig. 334
of larger wire than the lower part. It is often convenient to employ two different
sizes of wire in order to hit off more exactly the desired resistance. Assuming
that the resistance (cold) of the two parts of the coil will be 0-55 and 0-51 ohm
respectively, we get the size of square wire 0 -64 and 0 -58 cm. respectively. Although
the calculation is here given as a direct process (as indeed it might be if all sizes of
wire were available), in practice a little adjustment of the figures by trial and
error is required in order to make them fit the standard sizes of wire.

Calculation of the efficiency. The various losses in the machine are tabulated
in the left-hand bottom comer of the calculation sheet. In the case of a generator
direct connected to an engine, it is usual to include in the friction losses only the
losses in the outboard bearing. The amount to allow for friction and windage is
best obtained from measurements in similar machines. It is useful to plot the
results of tests in curves like those given in Fig. 222, so that we can quickly make
rough estimates of the friction and windage of rotating parts. The amount of
windage will depend very greatly on the shape of the rotating parts. Any pro-
jections which act as blowers will greatly increase it, so that some judgment must
be employed in using Fig. 222, which gives the friction and windage on rotating
fields of the ordinary sort which are not fitted with special blowers.

The iron loss at no load has been previously found to be 13,800 watts. Where
the performance specification is worded as Specification No. 1, Clause 16, the
no-load iron loss is to be taken in calculating the efficiency. The full-load iron
loss ia so difficult to measure, that the above method of giving the guarantees is
to be preferred. The field losses should include the rheostat losses, and therefore
are obtained by multiplying the field current by the exciting voltage. Adding
together the losses and taking the ratio of output to input, we get the efficiency
figures as given on the calculation sheet.

Wave-form of E.M.F. If we plot the field-form in the manner described on
page 14, we will find that the 5th harmonic (see page 22) is about -0-09 of the
fundamental. Referring now to Table XVI. page 313, we find that the value of /^
for an armature winding, having three slots per phase per pole, is 0*217 for the phase
pressure and 0-217 x cos 150°, or 0-187 for the terminal pressure. The value
of the 5th harmonic in the e.m.f. wave-form is therefore -0 09 xO 187 =0-0169.
Similarly, it will be found that the value of the 7th harmonic is 0-012. The 3rd
harmonic is, of course, zero, because of the star connection of the armature. As the
value of f^ is -966 x -866 = -835, the amplitude of the 5th harmonic will be
0-0169 -f -835 = -02 of the fundamental. The 7th is 0 012 -r -835 = 0144.

CHAPTER XIII.

ALTERNATING-CURRENT GENERATORS (cantinnedy

SLOW-SPEED ENGINE TYPE.

SPECIFICATION No. 2.

2180 K.V.A. THREE-PHASE GENERATOR TO BE DIRECT-

CONNECTED TO A GAS-ENGINE.

21 . The work covered by this specification is to be carried coSdiwons.
out in accordance with the General Conditions, a copy of
which is attached hereto and marked A.

«tc.

22. The work includes the supply, delivery, erection, etc., ^^^^ ^'

(See Clauses 2, 31, 51 and 80.)

Normal output
Power factor of load
Number of phases
Normal voltage
Voltage variation
Amperes per phase
Speed
Frequency
Regulation

Over load

2180 K.v.A. or 1760 K.w.

0-8.

6300.

6200 to 6500.

200.

125 revs, per minute.

50 cycles per second.

7 per cent rise with non-induc-
tive load thrown off, the
speed and excitation being
constant. 20 per cent, rise
w4th 0*8 power factor load
thrown off.

250 amperes at 6400 volts with
power factor between 0*8 and
unity.

Characteristics
of Generator.

334

DYNAMO-ELECTRIC MACHINERY

Nature of Load.

Flywheel type
or attached
flywheel

Shaft.

Parallel
running.

Exciting voltage 126.

Temperature rise afterl 46° C. by thermometer.
6 hours full load /55° C. by resistance.

Temperature rise after 1 65° C. by thermometer.
2 hours over load J 65° C. by resistance.

23. The generator is intended to run in parallel with five
generators of similar output and speed installed or to be
installed in the same power-house. These generators will
deliver a general electric supply to the town of ,
the load consisting partly of fighting and traction and partly
of induction motors in factories. The electric tramways in the
town are suppUed with continuous current at 650 volts, by
means of 50-cycle rotary converters fed from transformers
connected to the town 6300 volt supply. The generator must
be suitable in every way for this class of load.

24. The generator may either be of the flywheel type, or
it may have a flywheel rigidly attached to the revolving
field-magnet. In the latter case, the flywheel will be con-
sidered part of the generator, and must be included in the
price quoted. The construction of the flywheel and the
method of attachment must be indicated in the outline
supplied with the tender.

24a. The shaft will be supplied by the maker of the gas-
engine, and the Contractor shall 4 weeks before the date
fixed for delivery furnish the maker of the shaft with all
suitable gauges and information to enable him to turn the
shaft to the right diameter.

25. The Contractor shall be responsible for the provision
of a flywheel of the proper moment of inertia to enable the
generator to run in parallel with existing sets. The following
particulars are supplied to enable him to arrive at the best
dimensidns of flywheel :

(a) The gas-engine will have eight single-acting
cylinders working on an Otto cycle, there being four
impulses given by the engine per revolution.

(6) The speed is governed by controlling the amount
of gas and air admitted, and not by the hit-and-miss
method.

(c) A flywheel having a moment of inertia equal to
1500 tons at a foot radius will be sufficient to reduce the
angular irregularity to 1 in 250.

ALTERNATING-CURRENT GENERATORS 335

The Contractor may inspect the two generators and engines
already installed, and may take tachograph records at his
own expense. The gas-engine to be installed will be of the
same kind as those already installed, but no guarantee (other
than what may be contained in the above particulars) can be
given that it will operate in exactly the same manner as
the present engines.

The two generator sets at present installed run in parallel.
The interchange of power between the sets does not exceed
10 per cent, of the full load of one of them.

Heie may follow the Clauses Nos. 5, 6, 8 or its equivalent (see Clauses other clauses.
55 to 59, 60 and 273), 10, 11, 12, 13, 14, 15, 16, 17, 18, 19 and 20. In-
stead of 19 the following clauses may be inserted :

26. The following tests shall be carried out before the Teste before
generator leaves the Contractor's works : ^^^^ '

(a) As many coils as possible shall be inserted and Puncture test.
connected in the two halves of the stator frame, and the
whole shall be subjected to a test pressure of 13,000 volts
to earth for one minute.

(6) Any one coil may be chosen by the purchaser for cou tested to
testing to destruction. It shall withstand a puncture test '"*"^''*'"
of 3000 volts between successive turns and 16,000 volts
to earth applied for 5 seconds.

27. The following tests shall be carried out after the Test* after

. • i_ T 'i Erection.

generator is erected on site :

(c) The generator shall be run at full load, 0*8 power
factor, for six hours, and for two hours on the stated over-
load, and measurements shall be taken of the temperature
of the armature windings and iron and of the field windings
by thermometer, and of the field windings by resistance,

to see that the specified temperature rises above the sur- Temperatiure
rounding air are not exceeded. For the purpose of these ** '
tests the temperature of the engine-room shall be taken
three feet away from the generator, in a line with the shaft.

(d) Inamediately after the temperature run and while
the machine is still warm, an alternating pressure of 13,000
volts virtual shall be applied between the armature Puncture
winding and frame for ore minute, and an alternating
pressure of 1000 volts virtual between the field winding

and frame for one minute.

(e) A measurement shall be made of the exciting Exciting
current at 6300 volts at full load at unity power factor

and at 0*8 power factor.

336

DYNAMO-ELECTRIC MACHINERY

MngnetlzAtion

Short-circuit
Characteristic.

Regulation.

Parallel
Bunning.

Method of
determining
the efliciency.

(/) The generator shall be run at full speed at no load
with the field excited, and measurements shall be taken
to find the field current required at various voltages.

{g) The generator shall then be run with the armature
short circuited, and measurements taken to show the
relation between the field current and the armature
current.

(h) The regulation shall be determined by noting th6
current required at full load as prescribed in test (e), and
seeing what voltage corresponds to that exciting current
at no load, according to test (/). For this purpose 6300
volts shall be taken as the full-load voltage.

{%) The generator shall be synchronized and switched
in parallel with the bus-bars, while these are fed by the
two existing generator sets (which shall at the time be
running well m parallel between themselves). The new
generator shall run well in parallel on the bus-bars at all
loads and at any voltage between 6200 and 6500, whether
there shall be two or one of the existing sets running, and
whether these or either of them shall be loaded or un-
loaded. The new set shall not be deemed to run well in
parallel if a dead-beat wattmeter placed in circuit with it
shall show an interchange of power of more than 200 k.w.,
after due time has been allowed for any irregularity due
to switching to settle down.

(j) If any dispute shall arise as to the efficiency of the
generator, it shall be determined in the following manner :
The connecting rods shall be disconnected from the cranks
of the engine, and the generator shall be run as a syn-
chronous motor, being started up from rest with one of
the other generators in the station. When running at
full speed at 6600 volts unity power factor, the power
taken to drive it shall be measured by means of watt-
meters. The power so measured, after deducting 10 k.w.
for the loss caused by the shaft and cranks, shall be taken
as the iron loss, friction and windage. The PR loss in the
armature shall be calculated from the resistance of arma-
ture taken at 60° C. The excitation losses shall be taken
as the exciting current determined under (c), multiplied
by 125. The efficiency shall be calculated from the sepa-
rate losses foimd as above. The cost of making the iron
loss test shall be borne by the party calling for the test,
unless he can show that he was justified in doing so, in
which case it shall be borne by the other party.

ALTERNATING-CURRENT GENERATORS 337

THE DESIGN OF A 2180 K.V.A. THREE-PHRASE GENERATOR, TO BE

DRIVEN BY A GAS-ENGINE.

The principles which enter into the design of this machine are in general the
same as those which control the design of the smaller engine-type machine, but
the fixing of the fl3nvheel efiect in this case is a matter of considerable importance,
To arrive at the best flywheel effect to give to a generator under any given circum-
stances, it is necessary to consider shortly the laws which govern the parallel
running of synchronous machines.

PARALLEL RUNNING OF ALTERNATORS.

It is not within the province of this book to enter fully into the theory of the
parallel running of alternators. The matter is very fully dealt with in text-books
and in papers read before various institutions.*

We shall look into the matter with two main objects in view : (1) to- enquire
what information should be given by the man who is drawing up the specification
of an alternator which is intended to run in parallel with other machines, and (2) to
see what steps the designer should take to make sure that the alternator will run
well in parallel under the stated conditions.

For these purposes, it is well to remind the reader of the main principles
involved, and to collect the formulae to be used in a handy form.

Every synchronous alternator or motor when running in parallel with a net-
work is constrained to run in the true synchronous position by a moment which
behaves like the torque exerted by a spring ; that is to say, the turning moment
is proportional to the amount of displacement from the true synchronous position.
If any displacement from the synchronous position suddenly occurs due to any
outside disturbance, the field-magnet of the alternator swings about the central

*See Gisbert Kapp, ElektrotecJinische ZeitHchrift, vol. 20, p. 134 (1899) ; Goldsohmidt, ibid.^
vol. 23, p. 980, 1902; Hobart and Punga, Trans, Am. LE,E., vol. 23, p. 291, 1904; Punga,
Elektrotechniache Zeitschrifi, vol. 32, p. 385, 1911; Schiller, ibid., vol. 32, p. 1199, 1911;
Rezelman, LumUre Electrique, vol. 15, p. 67, 1911 ; Paper by E. Rosenberg, J(nir. Inst. Elec.
Eng.j vol. 42, p. 524; The Dynamo^ by Hawkins and Wallis, vol. 2, p. 998; Wechsdatroni-
Tnaschinen, by W. Petersen, pp. 248 (published by Enke, Stuttgart) ; A. R. Everest, " Some
Factors in the Parallel Operation of Alternators," Jour. Inst. E.E., vol. 50, p. 520, 1913.

The following articles are also of importance :

" Parallel Operation of Alternators," G. Benischke, Elektrotech. u. Maschinenbau, 25, p. 1009,
1907 ; L. Fleischmann, Elektrotech. u. Maschinenbau, 26, p. 329, 1908 ; H. CU^rges, Phys. Zeitschr.,
9. p. 265, 1908 ; O. Weisshaar, Elektrotech. u. Maschinenbau, 26, p. 555, 1908 ; G. H. Shepard.
Elec. World, 52, p. 271, 1909 ; " Ready Reckoner for Flywheel Meet in Armatures, etc.," H.
Luckin, Electrician, 62, p. 642, 1909 ; " Parallel Operation of Three-phase Generators with
their Neutrals Interconnected," G. J. Rhodes, Amer, I.E.E., Proc. 29, p. 639, 1910; J. R. Barr,
I.E.E. Joum., 47, p. 276, 1911; "Measurement of Relative Angular Displacement in
Synchronous Machines,*^ W. W. Firth, I.E.E. Joum., 46, p. 728, 1911 ; " Investigation of the
Swinging of Synchronous Motors,t^^eldmann & Nobel, ArchivJ. Elektrot., 1, p. 291, 1912 ; "Appa-
ratus for Measuring Irregularities in Speed of an Alternator, ^Boucherot, Soc. Int. Elect., BuU. 2,
Ser. 3, p. 557, 1912 ; " Influence of Torsional Oscillations of Shafts on Parallel Running of
Alternators," L. Fleischmann, Elektrotech. Zeitschr., 33, p. 610, 1912 ; " Bipolar Diagram of
Synchronous Alternators and Motors," Blondel, Comptes Rendus, 156, p. 545, 1913 ; " The
Synchronizing Couple of Synchronous Machines," Blondel, Comptes Rendus, 156, p. 680, 1913 ;
" Parallel Operation of Alternators with Composite Windings," Mossman, Elec. World, 61, p. 56,
1913 ; " Phase- Swinging of two Alternators coupled by Transformers," Gavand, Lumiere
Electr., 22, p. 103, 1913.

w.M. Y

338

DYNAMO-ELECTRIC MACHINERY

position like a pendulum until the energy of the swing has become dissipated.
The frequency of this phase swing we will call w,. This natural frequency of phase
swing depends upon factors which we will consider later. If the natural period
of the swing is the same as the period of a regularly recurring disturbance (such
as may be caused by the uneven turning moment of an engine), resonance is
set up, which may increase the swinging until parallel running is impossible. One
of the objects of the designer will be to avoid resonance.

Let us consider a two-pole machine running in parallel with mains of constant
alternating voltage and constant frequency. We can then conveniently take the
phase of the voltage of the mains (sometimes called the network) to be our phase
datum line, from which we can set ofi the angle of lag or lead of all other voltages
and currents.

Fig. 335 represents the armature of an alternator connected to the supply mains.
The arrow head on the circuit denotes the direction taken as positive for the

no. 885.

-Ao-]

FlO. 336.

purpose of the clock diagram (Fig. 336). When the field-magnet revolves, it gener-
ates an E.M.F. in the winding which is almost directly opposed to the E3f.F. of the
network. Thus the clock diagram in Fig. 336 would represent the state of affairs
where voltage OG makes an angle cr with the line of the network voltage ON. The
resultant voltage driving the current is given by 0/2 and the current by OC, Some
writers merely draw the triangle NOR to represent the state of afiairs, but this
is sufficient only if the sign of the various vectors in relation to the arrow head in
Fig. 335 is clearly ascertained.

As the synchronous reactance in the generator is usually very much greater
than the resistance, the current OC supplied to the mains lags about 90 degrees
behind the resultant OR, and is therefore almost in phase with OG and almost
180° out of phase with ON, Under these circumstances, the machine acts as a
generator, and there is a torque tending to slow it down.

If the field-magnet of the machine is behind 0N\ say in the position OM (Fig.
337), then the current lagging behind ORi will be nearly ISO*' out of phase with
OM, so that the machine will behave as a motor. That is to say, the torque will be

ALTERNATING-CURRENT GENERATORS 339

such as to tend to increase the speed. This torque, called here the s3aichronizing
torque, will be approximately proportional to the angle o-. In a two-pole nuichine,
this is the angle which the centre line of the field poles makes with the phase datum
line ON. In a machine having p pairs of poles, if a is the angular displacement
of the line of the poles, then

The angle <r is the displacement on the clock diagram which shows the electrical
relations, while a is the mechanical displacement. We will see later what features
in a machine determine the relation between the synchronizing current and the
^j^ngular displacement ; but. for the moment we will simply denote by lu the sya-
chronizing current per unit angle of displacement when the conditions are such
as to keep the power factor near unity. Then for any small displacement <r the
synchronizing current will be (r/„, and the synchronizing power will be <tIuE,
where E is the voltage of the network.

The synchronizing torque will be obtained by dividing this power, EIu<ry by the
speed expressed in radians per second. If Rpg is the speed of the generator in
revolutions per second, 2irRpg gives us the number of radians per second.

Thus the synchronizing torque in kilograms at a metre radius

- ^/«q- _0'01 62^/^0-

Let us suppose that we have a periodic disturbance, due, say, to the irregular
turning moment of the engine driving the generator, which follows the law

Qd sin 2imdt,

where Qd is measured in kilograms at 1 metre radius, and let us leave out of account
for the moment the synchronizing torque. The amount that the speed is changed
at each pulsation will depend upon the value of the flywheel effect ^mr^, and upon
the frequency of the disturbance rid- The increase in speed will follow the law :

1 9'SlQd

a = —

27r7irf Imr^

cos2'irndt, (2)

where ^mr^ is measured in kilograms at a metre radius*. The amount of the
angular displacement, a, of the rotor will be the integral of this, or

a^^^J:^sm2^ndt (3)

Thus we see that the displacement is directly out of phase with the disturbing
torque, and imder these circumstances any synchronizing torque will be added
to the disturbing torque. As the two torques are added, the phase swing will
be increased. The amount of the increase will depend upon the ratio of the syn-
chronizing torciae (brought into aci^ion by a displacement produced by a certain
disturbing torque) to the disturbing toraue.

Let us use the symbol q for this ratio. Then

Synchronizing torque Q«_

Disturbing torque producing it" Qd'^

340

DYNAMO-ELECTRIC MACHINERY

Consider first the case when q is less than unity. Then the final value of torque
will depend upon the value of the sum of an infinite series

l+g' + j24.^+^^ etc.

When gr is less than unity, the sum of this series is finite and has a value ^— — . That
is to say, the ratio of final oscillating torque to the initial disturbing torque is
= . Where q is less than unity, this expression gives positive values which

become greater and greater as q approaches unity, and infinity when q—\. For
values of q greater than unity the synchronizing torque is opposed to the disturbing
torque, and the greater the value of q the less the displacement. The reader is
referred to the very neat graphic constructions given by Dr. Rosenberg in his

0
c

■I-

a
C

Fig. S88.

ai* fo
h ^ to
c « 3ooe

(AJ S

S/rn^

FlO. 889.— Change of amplitude of damped oscillation with
change of frequency of the disturbance (see page 856).

paper, which are of great assistance in obtaining clear ideas of the relations of the
various quantities involved. As qQd^Qs, the ratio of the final synchronizing torque

to the initial synchronizing torque is j^- We will adopt the term " Wobble factor "

there proposed for the expression —2 —

The relation between original disturbing torque Oa (which is taken as 1 - gr)
and the final torque OA (which is taken as 1) can be seen from Fig. 338, which
refers to the case where q<l. Oc is the original displacement and 00 the final
displacement. Now we take our scales such that 00 represents the final synchroniz-
ing torque, and add it, Aa, to the original disturbing torque Oa, getting OA. We
see that if a total torque 1 produces a displacement which gives rise to a synchronizing

torque q, then the ratio of the total torque to the original disturbing torque is .

In the case of resonance we have q = l, and the oscillations go on increasing
until they are so great that the whole of the energy of the disturbance is absorbed
in the damping action of the poles. If the damping action is very small, the msu^hine
will go out of step before this point is reached. If the damping action is very
great, the machine may run in parallel, notwithstanding complete resonance. The
answer to the question how great the wobble factor may be before parallel running

ALTERNATING-CURRENT GENERATORS 341

becomes impossible, depends on the magnitude of the disturbing torque and the
eSectiveness of the damping action of the poles. As we change the frequency of the
disturbance, keeping the other conditions constant, the amplitude of the displace-
ment is gradually increased as we approach the frequency at which resonance occurs
in the manner indicated in Fig. 339. At the crest of the curve q = l, and the whole
of the energy of the disturbance is then expended in overcoming the damping
forces (see page 602). There will in most cases be various disturbing torques,
each with its own frequency. In a steam engine, even though there may be many
cylinders, there is usually a disturbing torque having the frequency of the revolu-
tions of the engine. Even though the amplitude of this torque may be smaller
than the amplitude of disturbing torques having higher frequencies, it may neverthe
less be the most important element to take into account, because the displacement
produced is inversely proportional to the square of the frequency of the disturbance
(see (3), page 339).

We have seen that the synchronising torque

^ 0-0162Jg/«<r

and that the maximum value of o-=»2r2 — 2~yr~r ^-P*
Therefore the maximum value of Qg during the swing is

Dividing by Qd we get

g^^^ =0-00403^ ^^^/-P , (5)

The critical value of flywheel effect which brings about resonance is the
value which makes ^ = 1. We have, therefore,

{I.mr^Ut. = 0-00403 -^* ^ ^-^ kilograms at a metre^ radius (6)

■lips X Ufl

Orjin British units :

(27n67'6%t. = 0-0000425^^^?^^ tons at a foot^ radius (7)

Or if we prefer to give the flywheel effect in kilograms on a metre diameter,

we have KL.x7)

GfZ>2^t. = 0-01612-^^^^^!^^ kilograms on a metre diameter (8)

-tips ^ ^d

Observe that in the above formulae EIu is in watts. If the synchronizing power
is expressed in kilowatts, then GD^ will be in 1000 kilograms on a metre diameter.

Let the ratio of /„ (see page 339) to the full-load current Ii be )8, so that
EIu—PEIi, We then get a simple expression for the critical flywheel effect per
K.V.A. of output, as follows :

G^^crit = ^^^p^^^^g^^ in 1000 kilograms on a metre diameter per K. v. a., ... (9)

Xtpg X Tiff

or, in British units,

{^mn^)cnt. = ^'^^^ ^ ^^^ in tons at a foot radius per K.V.A (10)

342

DYNAMO-ELECTRIC MACHINERY

As a first approximation, lu is sometimes taken as the current which flows in the
armature when the generator is short-circuited and run fully excited.* It has been

pointed outt that this does not give a very accurate
result, because on salient-pole machines the syn-
chronous impedance when running at unity power
factor is lower than when running on zero power
factor. The value of lu on a salient-pole machine
will in general be higher than Iq, the short-circuit
current.

As we have seen on page 295, the angle between
the centre line of the pole and the phase line of
the terminal voltage Et consists of two parts : one
part due to the lag of the terminal voltage behind
the generated voltage, denoted by f in Fig. 340,
and the other due to the di3tortion of the field,
denoted by <^ in Fig. 340. The angle ( can be
calculated from the ratio between the true armature
leakage flux and the working flux, as shown by the
example given on page 345. The angle <f> can be
calculated from the ratio of the armature ampere-
turns to the field ampere-turns in conjunction with the coefficient K^ given in

Table XVII. {

^, ,. ,. , , armature ampere-tuiTis per pole (all phases) ^

The distortion ancle <f> = ^-tj f —^ — z — a 4. ^^.i. ^ ^*

^ field ampere-turns per pole on gap and teeth ^

If we calculate C aiid <f> for full-load current Zj, then

^ 57-3

If we denote by p the ratio between the synchronizing power for cr = 1 and the

PlO. 840.

normal full-load power, then

)8 =

57-3
C+<l>'

Table XVII.

Ratio Polo-»«^.
pole-pitch

K^ in d^;reofl.

0-4

7-0

0-5

100

0-6

13 0

0-7

18 0

0-8

24 0

0-9

310

40-0

♦On a three-phase machine one must, of course, multiply the current per phase by 1'73 in
order to get the current /„, which when multiplied by <rE gives the synchronizing watts.

tSee references given on page 337. See also communication of Mr. Shu ttlewort h, Jotir.
Iiist. Elec. Engra., vol. 50, page 549.

*See •*Some Factors in Parallel Operation," A. R. Everest, Jour. Inst, Mec, Engrs,^ vol.
50, page 620.

ALTERNATING-CUREEKT GENERATORS 343

The excitation that is efiective in changing /3 is the excitation absorbed in
the air-gap and t«eth. This may change over a fairly wide range in the practical
operation of a generatoi, bo that there will be a fairly wide range in the value of
flywheel efiect that might cause resonance. For instance, for a certain generator
with lowest contemplated excitation, j8 might be 3, and with the highest excitation

For my other frequency or Bhort-clrcuit current, tbe Hywheel effect will be varied in direct

proportion to the Irequency or ihort-clreuit current,
I. Averse Bywheel elfect of a niu-en^nc having 1 imjiuUa per mvliiliofl for a cyclic irrcEularlty

>-,to. II. Ditto, i'^ilr-

it might be 3-8. Filling these two values in the foimula given above, we get two
values of the flywheel effect, and anywhere in between these two values there is
danger of resonance, for at some excitation q would be equal to 1. Furthermore,
even outside this range there are values of the flywheel efiect that may make the
wobble factor too great for satisfactory operation. In Dr. Rosenberg's paper referred
to above, some very useful curves are given in which the resonance zones and danger
zones for various types of engine and various speeds are plotted in handy form.

344 DYNAMO-ELECTRIC MACfflNERY

One of these referring to a gas engine we reproduce in Fig. 341. With a gas engine
working on an Otto cycle there is always danger of some disturbing torque having
a frequency of one-half of the frequency of revolution. Even if the gas engine
has many cylinders, so as to give several impulses per revolution, it may happen
through the setting of the valves that one of the cylinders gives a disturbing torque
every two revolutions. For this reason the upper curves in Fig. 341 have been
plotted by taking Ud only one-half of Rpg. The lower curves have been plotted
by taking na equal to Rpg. Other curves might be plotted for higher harmonics,
but one is not likely to be troubled with these, because the displacement of the
flywheel is inversely proportional to the square of the frequency of the disturbance.
It will be seen that the curves for the ordinary sizes of flywheel used with four-
cylinder gas engines lie mostly in between the two danger zones, but for some
speeds they are intersected by the danger zones. For these speeds it would be neces-
sary either to make a flywheel so big as to get completely above the upper curve,
or to increase the value of the short-circuit current so as to raise the curves which
mark out the danger zone, or to add a damper heavy enough to ensure steady
running, notwithstanding the resonance between the disturbance and the natural
period of swing of the alternator.

As these curves are only plotted for certain values of p and only take into
account the disturbances likely to be met with on the generator's own engine,
we must have recourse to the formula

^ 0'0l6l2xi8x»

^D^cTii. — jr~~ — 2^^ ^^^ kilograms at a metre diameter per K.V.A.,

or Imiiri^^t. = — ^ ^ ^^^ ^^^ ^ ^^^^ radius per K.V.A.,

■tipg X fid

when designing a flywheel to avoid resonance under circumstances not covered by
the curves.

In Purchaser's Specification No. 2 we have purposely made the conditions
rather diflicult to meet, in order to illustrate how to use the formula and adapt a
machine to meet difficult conditions.

METHOD OF FIXING ON SIZE OF FLYWHEEL REQUIRED FOR AN ALTER-
NATOR DRIVEN BY A PRIME MOVER OF IRREGULAR TURNING
MOMENT.

As an example, we will take the 2180 K.v.A. three-phase generator, particulars
of which are given on page 348. This generator is to be driven by a gas-engine
having four impulses per revolution, running at 125 R.P.M. under the conditions
stated on page 334.

The first step is to calculate the synchronizing power for a displacement of
the field-magnet of one radian behind the phase of the voltage of the network.
For this purpose we first find what displacement of the field-magnet would occur
at full-load unity power factor. The displacement, as we have seen on page 342,
consists of two parts : the angle BOC (Fig. 340) and the angle 4>, To calculate
DOC we must make a rough estimate of the armature leakage at full-load current.

ALTERNATING-CURRENT GENERATORS 345

By the method described on page 422, we find that the permeance of the stator
slot per cm. length of iron is 2-09. As the length of core is 34 cms., we have

2-09x34x2 = 143.

At a load of 200 amperes through the four conductors per slot, the total slot
leakage will therefore be

143 X 200 X 1-41 X 4 X 1 -257 = 204 x lO^.

To arrive at the leakage from the end windings, we take the coefficient ^"^=2-1
from Table XVIII. page 427.
The end leakage at full load :

200<^« = 2 -8 X (30 + 10) X 2400 = 2 -66 X 105.

Thus the full-load leakage amounts to 4-2 x 10*. This is 7 -5 per cent, of the working
flux per pole 5*96 x 10^ (see page 35). A leakage flux of 7-5 per cent, will make a
displacement angle BOC of

7-5

100

X 57-3 = 4°.

21-5
The angle 4> is found as follows : Ratio of pole arc to pole pitch is -5^ = 0*72,

but, there being a small bevel on the pole, the effective ratio may be taken at 0-69.
From Table XVII. page 342, K^ will be 17-5°. The effective armature ampere-turns
are 3150, and the ampere-turns on the air-gap and teeth amount to 5000. There-
fore <^ = 11°. Thus the whole angle of displacement on full-load current unity
power factor is 4° + 11° = 15°.

If 15° gives a torque corresponding to the K.W. rating of 2180, then a displacement
of one radian, or 57-3°, will give a torque 3-8 times as great. Thus /3 in formula
(10), page 341, =3 -8.

The next point to decide is, what is the natural period of oscillation at which
we should aim, in order to avoid resonance ? If the generator is driven by a gas-
engine, one of the frequencies at which resonance might occur is the frequency of
the camshaft, which runs at 62-5 R.P.M., giving nd = 104. We find that if we try
to make the natural frequency of oscillation of the generator n«, as low as 80 per
cent, of this^ we shall require a flywheel of enormous dimensions, which will be
very costly, and will greatly increase the friction of the bearings. Moreover, this
heavy flywheel would be very much greater than is necessary to reduce to a
workable amount the cyclic irregularity of a gas-engine having four impulses per
revolution.

We will therefore try whether a smaller flywheel, giving a natural frequency of

oscillation greater than 62-5 per minute, will do. We must remember that we must

not make the flywheel too small, or there will be danger of resonating with the

frequency of revolution 126 per minute. We will therefore aim at a natural frequency

of oscillation of about 90 per minute, so as to come well between 62-5 and 125.

90 per minute gives us n, = l-5. In formula (10), page 341, we have the required

flywheel effect

0-0425 X 3-8 X 24 . ^„ ^ . r . j-
^"2^08 — r^ — r^ =0*82 ton at a foot radius per K.V.A. ;

that is to say, 1800 tons at a foot radius for the generator in question.

346 DYNAMO-ELECTRIC MACHINERY

We may check this calculation from the formula given by Mr. Eveiest:

/o = 9-76^^

180x3-8x50

foot-tons

Taking /^ at 90, this gives us 4800 foot-tons of stored energy, at a speed
of 125 R.P.M. A flywheel effect of 1800 tons at a foot radius at a speed of 2*08
per second gives us : 1 1800 x 4:ir^ x 2082

2' 32-2 =4»00.

We have given the calculation here at length, in order that the reader may
understand the method. It would, of course, have been very much shorter to
refer at once to Dr. Rosenberg's curves given in Fig. 341. These curves refer to
a machine in which the synchronizing torque for one radian of displacement lies
between 3*5 and 4*2 times the synchronising torque at full load, and may there-
fore cover the case where j8=3-8. It will be seen from these curves that, if it is
desired to get completely above the curve 5^2 when the speed is 125 R.P.M., a fly-
wheel effect of 2*7 tons at a foot radius per K.V.A. will be required. This would

+6%
4-4%

^/^/V^/V^/^^/^w^/^7^/^\y'^v-^^^/^>Ay^vAy^v^V/^\yV\/^^^

PlO. 842.

call for a flywheel of 5900 tons at a foot radius. If, however, we content our-
selves with getting in between the curves F^ *^^ ^a* * flywheel of about 1800
tons at a foot radius will suffice. It will further be seen from the dotted curves
I and II that a flywheel of the size chosen is satisfactory from the engine-builder's
point of view.

In cases where there is any doubt as to the periodicity of the disturbing torque
or the amount of it, tachograph records should be taken of the engine under con-
sideration. Sometimes there is a source of disturbance which would otherwise
have been left out of account. For instance, if an engine drives a single-acting
condenser pump, this may have a serious effect upon the torque diagram. Fig.
342 shows a tachograph record taken on a four-cylinder compound steam-engine
running at 75 R.P.M. when running at three-quarter load with normal setting of the
valves. Notwithstanding the eight impulses per revolution, it is found in this case
that there is a very decided irregularity occurring once per revolution, which is
much greater than any of the other disturbing causes.

Damper or amortisseur. Where a generator is driven by a gas engine, and
in all cases where there is an irregular turning moment or an unsteady frequency,
in addition to selecting the best flywheel effect, we should provide the machine with
a suitable damper (see Fig. 346). Where the conductivity of the damper is high,
it^is possible to run synchronous machines in parallel notwithstanding resonance,
so long as the disturbing torque is not too great. When the flyivheel effect is such
as to give resonance, the whole of the energy of the disturbance is expended in

ALTERNATING-CURRENT GENERATORS 347

overcoming the forces set up by the damper, and the amount of the phase-swing
is just sufficient to call into being damping forces great enough to balance the forces
creating the disturbance. This matter is treated quantitatively on pages 352 and 601.

THE DESIGN OF THE 2180 K.V.A. GENERATOR TO MEET SPECIFICATION

No. 2 (Page 332).

Having fixed upon the approximate flywheel effect which should be given to
the revolving part in order to avoid resonance, we have to decide whether we wiU
put the whole of the flywheel effect into the magnet wheel itself, or whether we
will provide a separate fljrwheel. It will be found that for slow-speed generators
of output less than 3000 K.W. it is more economical to provide a separate flywheel.
The economical diameter for the generator is generally too small to give it much
moment of inertia, and therefore if we try to get the desired flywheel effect into
the magnet wheel, we must either add an enormous weight to a magnet of reason-
able diameter, or we must increase the diameter so much as to greatly add to the
cost of the armature surroimding it.

We will, therefore, in this case fix upon the diameter of the field-magnet from
the considerations which are set out on page 299, and add a flywheel to give the
desired moment of inertia. It must be remembered that in any case the magnet
wheel will be too great to be shipped in one piece, so that it will have to be con-
structed in parts connected together by links or bolts.

If we had a frame with a bore of stator of 464 cms., that would be quite suitable ;
but the diameter might be changed over a fair range without appreciably affecting
the cost of the generator. To get the approximate length we might take the D^l
constant at about 4 x 10^. This gives us I about 32-5 cms.

It is found that with generators of this diameter there is very seldom any trouble
in meeting the regulation guarantees. The air-gap must be made of reasonable
length to avoid " pulling over," due to accidental unbalanced magnetic pull, and
if we have a reasonably great flux-density in the gap, so as to use our material
well, it will be found that the ampere-turns per pole are sufficient to give us even
better regulation than that called for in Specification No. 2.

One, therefore, begins the calculation of a machine of this size with a rough
calculation of the magnetic pull. We would like a flux-density in the gap of about
9000 c.G.S. lines per sq. cm. There will be 48 poles to give 60 cycles at 125 B.PJf .
Allowing a pole arc 0-64 of the pole pitch on the diameter chosen, each pole will
have an area of about 600 sq. cms. It would not be well on a machine of this size
to have an unbalanced pull greater than 5000 kilograms actual, or say 15,000
kilograms, neglecting saturation, for one milHmetre displacement ; so, from the
formula on the top of page 60, we get :

405 X 10-8 X 90002 X 48 X 600

^'"'""'•^ i5;ooo '

^ = 6*3 mm.
Let us take g at 0-65 cm. and B at 9000, and see about how many ampere-
turns we will want on the air-gap :

0-65.X 9000x0-796 = 4650 a.t.
This, as we shall see later, is sufficient to give us the required regulation.

348

DYNAMO-ELECTRIC MACHINERY

DiUeyj?»/iW/?je.. .19/3 .- Type <?-^t AJ^.s

KV A.;?/:^.; P.Rr.«.; PhMe.'^...; Volte .^^(^
4*rft....«^»-..Amp« p. zvckA.ZQQ, Amps p. br.

Cn«omer.<V'^X.<?4^.^'y<?W<r.<:A; Order No

; Amps per tecJUQQ. ; C7cles.<l^(2[.^..;-R.P.lf/?!$ ; Rotor AmpOt^ ;

Temp, rise .^.f.C-. fttgii\a!6tm.Qjp...SJ:,TJ^.OwAi»Aj^:^3kSSf^

Quot No» » ; Pcrf. Spec.

... Fly.wfaeel effccu/6[<?.<?^xnL

^~~*^^Circ«m.^4«^.; Gap Area^^<X?.; ^^
Air «— — . AgB

A(.B .w^..; poos. laZa »- ;UZa D*t,xRPi| ^^^ ^^^

.4.-.<?..^.^C?.f._.j UZ.^^,^.QOQ ;cT5r«..-23^. K.V.A^ztf.'.':^:

^K, .*4 ;,..igg<?g.Volts^:.4. X igriOgx /.7J?.<tu.4.-g!

Arm. A-T. p..pole.i?/.5^

.-Max. Fld..A.T. /^,/7.(7^?

Armature.

Stat.

o
o

I-

3
"D
C

o
o

Dia. Outs..
Dia. Ins

Gross Lei^t^
Air Vents

Opening Min Mean

Ak Velocity

364

A±^

>^Q65^

^Simper sjs<:j

Net Lengthi2i2L75x 89

Depth b. Slots

Section 3QQ. Vol.

Flux Density.

Los.slCtf^p.cu.gg?-Total
Buried' Cu.£<20l2Tota]
Gap \ttii4:9AQCL- Wts
Vent Area2ig2i2gg Wts
Outs. Area 62^ Wts

_ //

'&QO0 5X000.
21.000

ZQ^tQO

13M0

No of Segs
No of Slots
K.

\^2

Mn.Ciic.

Section Teeth .
Volume Teeth.
Flux Density.

f 25, OOP

JjossO'J&p.cxi £^?Z..Total

/%200
\2CkPOQ_

Weight of Iron -

Star ot Meah Throw

Cond. p Slot

Total Conds

Size of Cond. 1^5 X
Amp. p. sq. CTTl»

•65

Length in Slots^ 34_
Length outside ..^Osum
Total Length

920

25,200

44 70

54.700

</-2^

im-

I'l2',2'lt^ 3'tO

1723 I

0^52 sq.cm.
385

Wt. of l.cSo^^^Total
Res. p. 1,000 *.<d2fiTotal

Watts p.zzti

Surface p. CO-

Watts p. Sq.iaa

' ggx '057

•00/2

750

'64
63

Jipo

057

f2^C

^3Z Slots

ffr

<---S'J7"»

A ; r

?-:

Y V

f^-zai

TK'TK

; . jfSFb/es

C— 34 -M
...^..Vents

-- tS'S—^

Field

Rotor.

Dia.

\ Total Air Gap

Gap Co-eff. K,

Pole Pitch 3g? Pole
Kr

Arc

Flux per Pole

Leakage n.l /^ f.l j^"f
Area4<S^ Flux density _
Unbalanced Pull

No.ofSeg
No.of Slots
Vents

Mn.Circ.
X «

K, Section

4S2'7

'63

/'04

/9'5

'65

6'23X/0

3-63 x/O

ia.ooo

6,SQQ kUagnx

Weight of Iron^g/g^ 1 4/<70 kftog.

AT p Pole n. Load
A.T.p.Polef.Load

Surface ^^

Surface p. WattS-

V R

I R

Amps

No. of Turns.

Mean 1. Turn ,
Total Length .

Resistance

Res. pen. 000.
Size of Cond.

Conds. per Slot

Total

Length

Shunt.

9500

8oi*ios« Comm

say to.

000

^6.600
9400 f920
~2I'6K.W.

-P6
222

iqs

2/60 rri

0-3 X 3-5

Wt. per i,ooojn_
Total Wt.

Watts per Sq..^:^
Star or Mesh _
Paths in parallel

890.
114

4^0QO}^,QQOcor€

_'3d qh/n^ cold -4'^ hot

^ fsg. cm.

-I-

Magnetlzatlon Curve.

Core

Stator Teeth
Rotor Teeth
Gap

zsjion.

Pole Body JLLi
Volie

Pole Body N,L

Scotlon. Lsngth

<^S,506 -65 ^I50\

480 a (6800' 30

^SQQ,\/oiis.

Ll.ytyjk.l.

44O0
240

r*TS&-:

.^3C<>V0!t8.

kT.P*iHAT.

Ql^jA^.. iy,^'>o\ ya ^40 iBZSP,

BQ50
17.100

25&J

4S0O
560

\5700
240

EFFICIENCY.

Friction and W..

Iron Loss

Field Loss

Arm. &c. TR

Brush Loss

1725

21-6 19 17

foa-o

Output
Input

WMaso

Efiiciency %

\230£,

J-^MSS

lasa

1310

i4Q4

93-4

JA IS

£ hi

325.

77-5

433

959 \5I5

SI 2

3380'

6.600.Vo\ls.

j A.T.pxi»j A.T.

9300
f8^

^qJo.

920
\657Q_

4-CO

^sa

Commutator.

Dia

Bars

Volts p. Bar.
Brs. p. Arm .
Size of Brs.
Amps p. sq. .
Brush Loss _
Watts p. Sq. -

Speed

Mag. Cur ' Loss Cur.
Perm. Stat. Slot 2' I

„ Rot. Slot X =

Zig-zag _.^

2 x34 'X2'l ^143^
177x200x4 X 143 ^2-04
End 2-1 yi42'5x24aa^ 2' 1 6

Amps : To\..4-2X/0^

:X. =

Imp. V +
Sh. cir. Cur —

Starting Torque
Max. Torque _
Max. H.P

Slip

Power Factor

Leakage -4^ -^ .i^Ip:

ALTERNATING-CURRENT GENERATORS 349

We have a provisional figure for Agy namely 464 xttx 32-5, and this gives us
-4^8 = 4-25x108, from which we get the approximate Za = 1870. Now we would
like the number of conductors to be divisible by 3 x 48. As 12 x 144 = 1728, we
choose that number of conductors and work back to the correct AgB, This comes
out 4-6 X 10^, as shown on the calculation sheet, page 348. Drawings of the
generator are given in Figs. 343, 344, 345 and 346.

The ampere- wires ZaZa = 346,000, and the armature a.t. per pole =3150.
Add to this 7*5 per cent, of 4650 (see pages 283 and 345) to get the approximate
A.T. per pole on short circuit, 3500. From an assumed saturation curve, and by
the aid of Fig. 312, we can judge with fair accuracy that the regulation will be well
within 8 per cent, on unity power factor. Thus the above preliminary data are good
enough on which to found the design.

It is imnecessary to go in detail through the calculation sheet, as the general
method of working is the same as that indicated on pages 316 to 332 in connection
with the 600 K.w. generator.

With our increased AgS the length of armature iron comes out at 34 cms.,
in order to keep the density in the teeth at 18,200. The total losses to be dissipated
by the iron surfaces of the stator come out at 53,000 watts, and the watts which
the frame can get rid of with a temperature rise of 45® C. come out, on a conser-
vative estimate, at 54,700, so we expect to meet the temperature guarantee in
that respect.

The size of the armature conductor is settled by the considerations given on
page 323, and in this case a current density of 385 amperes per sq. cm. of copper
would appear to give us a temperature rise of 12** C. above the surrounding iron.

It will be seen from the drawing that we choose a parallel pole body in this
case. The number of poles being great, the angle between the neutral planes
bounding a pole is very small, so that a parallel pole body and a field coil with parallel
sides fill the space fairly well, leaving just nice room for ventilation.

The magnetization curve is worked out as indicated in the calculation sheet.
The figures are given for ampere-turns on the pole body with the degree of satura-
tion brought about by the leakage at full load (from which cAn be plotted the
increase-due-to-leakage curve, see p. 331), and also for the ampere-turns on the
pole body with the smaller degree of saturation brought about by the leakage
at no load. From the latter figures we get the no-load magnetization curve.

It may be of interest to work out in detail the cooling conditions of the field coils
of this machine, as a further example of the use of the formulae given on page 233.

The cooling coefficient for the ends of the coils :

he =00011 X (1 + 1 -2 X 125 X 11-2 x 221 x lO-^),

A, =0-00505.

We are allowed 45° C. rise of the field coils above the air. The coils are of copper
strap on edge, so that the heat conductivity of the coil is very good. We are there-
fore justified in taking the temperature difference between the surface of the coil
and the air at 40° C. when running. The total surface of the ends works out at
46,600 sq. cms. Therefore the heat dissipated by the ends is

40 X 46,600 X 0-00505 = 9400 watts.

350

DYNAMO-ELECTRIC MACHINERY

Fio. S48. — 2180 K.v.A. S-phase 50-cycIe generator and flywheel, designed to be direct connected

to a gas-engine running at 125 b.p.1[.

ALTERNATING-CURRENT GENERATORS

351

2Sd§cimetre$ 20

J-1

7fOBt

I I

O 10 coihrnetm

' ■ ■' ' ' i' ' ' / I ■ I ' ' |.M I

^ Z I O inches tZ

Fig. 844. — Showing flywheel consiBting of eight sectoTS of cast steel. The centrifugal force
on each sector is borne by the arms, which are bolted together and secured by shrink rings on
the central boss.

352 DYNAMO-ELECTRIC MACHINERY

The cooling coefficient for the part of the coil between the poles is

Ai = 1-5 X 10-8 X 125 X 11-2 X 221 xO-21.

Here s = 1 -3 and Z = 28, so that J? =0-21,

Ai =0-001.

The total surface of coils between poles is 48,000, so that the total heat dissipated
from this surface is

40x48,000x0-001 = 1920.

It is interesting to note that this is less than one quarter of the heat lost from the
ends, notwithstanding the larger cooling surface.

The thickness of the insulation around the pole body is 0-25 cm., and the heat
conductivity may be taken at 0-0012, allowing for some air-spaces. A pole of this
kind will not rise in temperature more than 10° C. above the air when running,
so we may take the mean difference in temperature between the coil and the pole
as 30° C.

Watts per sq. cm. = — ^ ^ =0-14.

The total surface may be taken as 95,000 sq. cms.

95,000 x014 = 13,300 watts conducted to the pole.

So that the total watts dissipated for 45° C. rise are 24,620. This is a con-
servative estimate, as the coefficients given in the formula are " safe " coefficients.
Let us now, as a matter of interest, calculate the watts dissipated in the rough
manner described on page 232. If we allow 1 sq. in. per watt, or 0 155 watt

per sq. cm., we have

0-155x189,600 = 29,400 watts,

a result which is probably not far from the mark. A calculation of the actual
watts lost in the coil at full load (0 8 power factor) gives us 21-6 K.W., so that the
coil will be safely below 45° C. as measured by thermometer.

The figures given for the stator leakage in the bottom right-hand comer of
the calculation sheet are explained on page 331.

Calculation of the effect of the damper. The simplest method of expressing
the effectiveness of the damper is to regard it as the squirrel cage of the rotor of
an induction motor, and to find the slip at full load which the machine would have
when run as induction motor. At full load the ampere-wires on the rotor must
be equal to the working ampere-wires in the stator. Now the working ampere-
wires in the stator, when carrying a load of 1750 k.w., are 278,000. Let ua take
the damper shown in Fig. 346, consisting of three copper rods through the pole,
and one rod on each side, making five rods per pole. Each of these has a cross-
section of 2-6 sq. cm. There being 240 bars in all, we would have a virtual current

278 000
per bar of <^^— = 1160 amperes when running as an induction motor. The

current in the end connections would be

i-x '— ^ X 0*637 = 1850 amperes virtual.
48x2 ^

ALTERNATING-CURRENT GENERATORS

PlO. 345.~3»ction ot 2180 K.

364 DYNAMO-ELECTRIC MACfflNERY

Now the resistance of all the rods in series, allowing for joints, is about 0*006
ohm, so that the loss in all the bars would be 1160x1160x0006=8100 watts.
It is impossible to calculate exactly the resistance of the end connectors (or end rings),
because the resistance of the joints is such an uncertain quantity. If the joints
are well made, one may allow for them by adding 100 per cent, to the calculated
resistance. The total length of copper in the two end rings is 29 metres, and it has
an average cross-section of 9-6 sq. cm. The resistance of the whole in one length,
without joints, would be 0-00051 ohnL Take the resistance with joints at 0-001
ohm. Then, as the virtual current flowing in these end rings is 1850 amperes,
the loss in them is 3500 watts, giving a total loss of 11,600. Now the slip of an
induction motor is equal to the ratio of the I*R loss on the rotor to the total power
supplied to the rotor (see page 433). Therefore the slip at full load with this damper
acting as the squirrel cage of an induction motor will be

^^=00066 (or 0-66 per cent.).

Denote this slip by 8,

EI X 1 *73
The full load torque is --— — kilograms at a metre (see page 339).

9*81 X ziirRpft

This torque is obtained with a relative angular speed between rotor and the

revolving stator field of -^ , where n is the frequency and p the number of pairs

T
of poles. Therefore, for an angular speed of 1 radian per second the torque will be

^^.^{^^'!?^^^ kUograms at a metre.

9-81 X 2ir X IJp, X 27rrw *

If a is the relative angular velocity between the rotor and the revolving field
of the stator, the torque due to the damper at any instant is

a — — — - -:^ — i-- kilograms at a metre.

9-81 X 2^ X /^p, X 2^ns ^

We make use of this formula on page 601, where the general theory of phase-
swinging under the influence of a damper is considered.

It is convenient to speak of a damper as a 1 per cent, damper or a 2 per cent,
damper, according as the slip at full load would be 1 per cent, or 2 per cent. The
damper worked out above would be described as a 0-66 per cent, damper.

It is interesting to enquire how far a damper such as the one illustrated in
Fig. 346 would be effective in preventing excessive phase-swinging in the event
of the disturbance being such as to cause resonance.

It will be seen from the theory given on page 602 and the example worked
out on page 356, that the amplitude of the phase-swing when resonance
occurs is such that the disturbing force is exactly balanC'Cd by the force exerted
by the damper. The amplitude of the phase-swing is proportional to the dis-
turbing force, and inversely proportional to the conductivity of the damper and
the frequency n<i.

Where a damper has sufficient conductivity, it may reduce the phase-swinging
to an amount which makes running quite possible even though ^ = 1. The amount

Fig. 846. — Blevation partly in section. The damper consists of three copper rods and a
copper washer around the pole. The dampers are inter-connected by copper links.

356 DYNAMO-ELECTRIC MACfflNERY

of the phase-swinging produced by a given disturbing torque can be approximately
calculated by the method* given on page 601.

Example 47. In the generator worked out on page 348, having a damper like that described
(0*66 per oent.)» and a flywheel effect of 1*7 x 10^ kilograms at a metTe^ find the amplitude of
the phase -swing a when the disturbing torque in kilograms at a metre follows the law
8500 sin 2ir x 1-5 x f. As we have here 71^ = 1-5, q=l, so that a would be infinite, if it were not
for the operation of the damper.

85008in( w/+^
a= -

nn ( (

w=2irx 1-6 =9-42,

,_ 1750xl0«x24

9-81 X 2-08 X 6-28 x 6-28 x 50 x 00066" ^ '

w6 = l-49xI0»,
8500

a= -

Bin ( tot +^\= -5-7 X 10~'sinf w^ + ^ y

1 -49 X 108

That is to say, the displacement lags 90" behind the disturbing torque, and has a maximum
value of 0*0067 radian, equivalent on a two-pole machine to 0*0057 x 24=0*137 radian, or 7*8

electrical degrees.

1*7 x lO''
In this case a = —^^Q^ — ; 6 = 1*58 x 10" and c = l'55xl(^. We therefore have a«'=c;

that is to say, the forces required for the acceleration of the flywheel are just supplied by the
synchronizing forces, leaving 6d= 8500 sin arf.

EFFECT OF HIGHER SPEED ON THE DESIGN.

If the speed of the generator were higher, say 150 R.P.M., and we wished to
keep the same peripheral speed, we should have to reduce the diameter. Now,
the output changes approximately as the square of the diameter, so the outputs
of machines of the same peripheral speed and the same length will vary approxi-
mately as the diameter. In this case, if the speed specified had been 150 R.P.M.,
we should have been compelled either to run at a higher peripheral speed or to
lengthen the frame. The best plan would be to choose a diameter of about
410 cms. and a length of 36 cms.

The way in which the number of conductors is increased on a machine of smaller
output, but of the same voltage, is illustrated in the calculation sheet of the 1800
K.v.A. generator given on page 357. Here the speed is 150. The number of poles
is 40, and the diameter has been reduced from that of the last machine in ratio
of the number of poles. The total AgB is reduced in the same ratio, and conse-
quently the conductors have been increased from four per slot to five per slot.

* In the case of a generator, running in parallel with a network of constant frequency, and
driven by an engine which exert-s a disturbing torque, Qa sin 2vndtf the equation of motion is

aa + 6d + ca = Qa sin 2Tmdt,

where a, b and c have the values given on page 601. Writing 2irnd = ^ and (aw* - c) = k, we have

When 7 = 1, A;=0, then a= -

= r- - -*^-^= sin I (at + tan~^ -y- ] .

Qaam(u>t+~j

lab

ALTERNATING-CURRENT GENERATORS

357

Dau.^.t/it'/y....iQ f2. Type §.'T'. ^.-.Gi..

KVA./e.<>!C?..; P.F.rA.; PhweJ. .; Volt«.ft?<»."r.
Jkmp* p coai./SQ. iwpi III t>

Fi
Air

I-

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,<^?^; Ainpe per ter../.?.?. ; CjcU».SlO.. .,.; R.P.lI./^.<?....; nnii Aiuue

.iS>:/»&.T«iBp. riM.5iC7.?.C ReBul*t«m.j»..%..^>^fi/ Owlimd ...^.'^.yp..

Order No.

lZI.Z..xQ^hx9MfiOQ^

AgB

; Qoot. No. ..; Perf. Spec ; Fly-wheel effect

hm e.i?.:g.y/g* : POM. J« «« ^<>0X>O9...\ l« «« D' L X B P jT

ICV.A.

'4-'l*iO'cm^

nu .:.*. J .jSf^f^. Voit*-i.:.4. x ^;?i5. x ^o/^^ 4.* !$.<?..

-; Ann. A.T. p. piM,.i3.4-O0..

FfcL K.lJhQQQ/fP:fi9l!f

Armature.

Dia. Outs

Dia. Ins —

GroM Length

Air Vents

Opening Min. _ Mean

im^'^iAaaecf aLattstc.

Net LengthJ2£.-dx-89
Depth b. Slots
Section ^gg

Flux Density.
Losa:fi£.p.GiLiL£ZLTotaI
Buried Ol J:fifi2TotaI
Gap knaJtAJiOO. Wts

Vent Area^&i202.; Wts
Outs. Areae&iffiC: Wts

NoofS^gs
No of Slots
K.

Section Teeth
Volume Teeth
Flux Density.
Loss ^p.cu

Weight of Iron

Length m Slots-ljSl

Surface p. Jit^'^f

Watts p. Sq.^/^

0'32

Field

Rotor.

Dia.

I Total Air Gap

Gap Co-eff. K,

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jafca

Flux per VtA^JLJ^&jUOl.
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No.ofSeg
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O^SL

JUL

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S^tOSL

A.T.pPolen.Load
AT.p.Polef.Load
Surface

Star

Cond. p. Slot

Total Conds

Size of Cond. -.££.xlfi£
Amp. p. sq.iC/^-

Surface p. Watt

I. R.

Amps. ___

No. of Tunis

Mean 1. Turn

Total Length

Resistance

Res. per x,

Length outside jJSSSnxm
Total Length

Size of Cond

Conds. per Slot.
Total

Wt of i.ooo-i2£2_TotaI
Res. p. 1,000 '^^Total
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Length

Wt. per z.ooo-aiu
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Watts per Sq.C^

Star or Mesh

Paths in parallel

^'^0

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Core

Stator Teeth
Rotor Teeth
Gap

19.900 4- '8

Pole Body CxLi.
Yoke

Po/eBodyAf.l.

Section.

26S

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9670 ♦

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360

irflCICNCY.

Friction and W.

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Field Loss

Arm. Ac. PR_
Bnish Lo» —

Output

Input ■ ■■
Efficiency -5s_

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Bar.

Dia.

Bars .
Volts p
Brs. p. Arm
Size of Brs.
Amps p. sq.
Brush Loss
Watts p. Sq.

.5peed

Mag. Cur.
Perm. Stat. Sk>t
„ Rot.Sk>tx
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Loss Cur.

2 X

177
Eod

Ns,

X
X X

Amps ; Tot.
; X. «

« +

Imp. y/ +

Sh. cir. Cur

Starting Torque
Max. Torque ^

Max. H.P

Slip

Power Factor

358

DYNAMO-ELECTRIC MACfflNERY

This enables the length to be slightly reduced. The general method of working
out the machine is as before. In this case the pole body has been made 16*5 cms.

12000

uooo

10000
$000

oooo
jooo

^ 6000

•s

5000

, ,4000
^3000

2000

1000

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m 3c

>00 4i

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00 6C

yoo n

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}00 JOt

w m

700 JH

100

Ampere-Turns per pole

Fig. 347. — Magnetization cnrve of 1800 k.v.a. generator, showing method of finding effect of

saturation on magnetic pull.

wide and of 28 cms. axial length, instead of 18 cms. wide and 27 cms. axial length.
The effect is to give a little more cooling space between the coils, and a little more
saturation in the poles.

The no-load magnetization curve and the increase-due-to-leakage curve of this
machine are given in Fig. 347. The method of working out the effect of the satura-
tion on the imbalanced magnetic pull is given on page 60.

CHAPTEE XIV.

ALTERNATING-CURRENT GENERATORS (caniinued).

WATER-TURBINE TYPE.

SPECIFICATION No. 4.

2600 KV.A. THREE-PHASE GENERATOR TO BE DRIVEN BY A

WATER TURBINE.

31. The Contractor shall supply and erect at the power- sztent oc

house of the Purchaser, situated at

tor having the characteristics specified below.

y a genera-

Work.

32. Normal output

Power factor of load
Number of phases
Normal voltage
Voltage variation
Amperes per phase
Speed
Frequency
Regulation

Over load
Exciting voltage

2500 K.w.^at unitypower factor. Ghanusteriitict

2500 K.V.A. at 0-8 power factor. ^^'«^*"-

Between unity and 0'8.

6900.

6800 to 7000.

210.

600 revs, per minute.

50 cycles per second.

12 per cent, rise with non-
inductive load thrown off, the
speed and excitation being
constant.

18 per cent, rise with 0*8 power
fector load thrown off, the
speed and excitation being
constant.

263 amperes at 6900 volts power
factor Between 0*9 and unity.

90 volts.

* Where the jpower factor of the load will probably be near unity, it is beat to call for the
full K.V.A. at unity power factor and a redaoed k.w. at lower power factors.

360

DYNAMO-ELECTRIC MACHINERY

Horisontal
Shaft.

Coupling.

Shaft.

FooDdations.

Temperature rise after] 45° C. by thermometer.
6 hours full load j 55° C. by resistance.

Temperature rise after ] 55° C. by thermometer.
2 hours over load /65° C. by resistance.

33. The generator must run on horizontal bearings, and
be designed to be directly coupled to a water turbine with
horizontal shaft.

34. Both halves of the coupling will be supplied by the
makers of the turbine, who shall be responsible for the proper
working of the coupling.

35. The shaft and other parts of the generator shall be
strong enough to withstand the shocks which may come upon
it if the generator is short circuited at full voltage.

36. The foundations for the generator will be supplied by
the Purchaser, but the Contractor shall supply the holding-
down bolts and plates, and be responsible for the erection on
the foundations and the grouting in of the bedplate.

37. Within ten weeks of the receipt of the order, the
Contractor shall supply sufficient particulars of the bedplate
to enable the foundations to be laid out, and at a convenient
time shall supply the foundation bolts and plates and a tem-
plate for setting out the same. He shall also supply within
ten weeks of receipt of the order sufficient particulars of the
shaft to enable the coupling to be manufactured.

to?So v^ow* 38. The revolving part of the generator shall be designed
Speed. " to run with safety at a speed of 1080 r.p.m., so that in the

event of the water turbine running away no serious accident

shall happen. The generator shall be run at 1080 r.p.m.

for ten minutes at the Contractor's works before being

despatched.

39. The generator is to form one of a number of similar
machines, each coupled to its own turbine and electrically
connected in parallel on the same bus-bars. The load will
consist of a general electric supply to various towns and
villages lying within five miles of the power-house and fed
by overhead lines. The generator shall be suitable in every
way for this work.

HouM.' ^^^" *^- ^ P^^^ ^f ^^^ power-house accompanies this specifica-
tion, showing the positions of the proposed water turbines

Paitlciilan for
FonndatioDS
and CoupUnff.

Ruoning
GonditioDfl.

ALTERNATING-CURRENT GENERATORS 361

and generators and the space available for the same. The
general method of carrying the weight of the machinery is
indicated.

41 . The Contractor shall supply with the generator a bed- Bedplate.
plate or sole plates suitable for the proposed foundations.
There shall be two self-aligning bearings fitted with automatic
oiling arrangements and means of adjustment.

42. The cables from the terminals to the switchboard will cawea.
be provided by the Purchaser.

43. The generator shall be star-connected, and the centre star Point.
point of the star shall be brought to a terminal which shall

be clearly marked " star point."

44. The ends of the three-phase winding may be brought Terminaia.
out from the armature by means of cables provided with
suitable sleeves for connecting to the Purchaser's cables.

Such terminal cables shall be insulated with waterproof
flexible insulation material of high quality, which shall not
be rubber or any material liable to be softened by heat.

Here may follow the following clauses, or such of them as are suitable
under the circumstances of the case : Clauses Nos. 8, 9, 10, 12, 13, 14, 15,
16, 17, 18, 19, 20, 23, 26, 27, 60, 61, 69, 73, 74.

DESIGN OF A 2500 K.V.A. WATER-DRIVEN, THREE-PHASE GENERATOR

TO MEET SPECIFICATION No. 4.

The main difficulty in the design of generators of high output and high speed,
such as are required for connecting to water turbines, lies in providing a sufficient
factor of safety in the mechanical design. It is usual to provide a fair factor of
safety at a speed 80 per cent, higher than the running speed, and as this is usually
already high, a special construction of the field-magnet is required to resist the
great centrifugal forces. The centrifugal forces are not in general as high as in
some turbo-generators ; but as the number of poles on water-driven generators
is usually great enough to make the use of salient poles economical, one commonly
finds on these machines a type of construction peculiar to them. The pole pieces
are very often made separate from the field spider, so as to admit of overhanging
pole spurs to support the field coils. Where the pole pieces are held on by bolts
it is usually necessary to provide a very large number of bolts for each pole, so that
a large percentage of the pole area consists of the cross-section of bolts. This is
rather expensive. A cheaper construction is to provide a very large dovetail or
two dovetails at the root of each pole. Another construction is that shown in Figs.
348 and 349, in which portions of the polar extensions are interleaved with por-
tions of the spider, and provided with cotter pins or bolts running axially. This

362

DYNAMO-ELECTRIC MACHINERY

construction makes good provision for resisting the centrifugal forces, and allows
the field coils to be put under considerable pressure.

The design which we have taken to meet Specification No. 4, is one by the
Oerlikon Company. It is illustrated in Figs. 348 and 349. The calculation sheet
is given on page 364.

It is unnecessary to go through this calculation sheet in detail, as the general
method of design will be imderstood from the description of the 600 E.w. on pages

Figs. 848 and 840. — 2500 K.y.A. S-phase generator, 7000 volts, 50 cycles, designed to be coupled to

316 to 332. The diameter is usually fixed by taking the largest diameter which
can be built economically with a sufficient factor of safety. If too small a diameter
were chosen, the axial length of these machines of great output would be too great
and the cooling conditions would be bad. In this case the diameter is 160 cms.,
and with a D^l constant of 4-6 x 10* the axial length comes out at 75 cms. The
peripheral speed is 50 metres per second, or just about 10,000 feet per minute
— a suitable speed with this construction of rotor.

The tips of the pole are made to well overlap the coils, so as to give good mechani-

ALTERNATING-CURRENT GENERATORS

363

cal support, and thus the ratio of pole arc to pole pitch is rather high, namely
0*71. As there is very little bevel on the pole, the electromotive coefficient Ke
comes out as high as 043. In other respects the working out of the machine follows
very closely the method given in the previous examples.

As this machine has salient poles, it is of interest to work out the ampere-
turns required at full load 0*8 power factor by the method described on page 294,
and to compare the result with that obtained by the method described on

a water turbine nurning at 000 revs, per minate. See Speciflcation No. 4 and Calculation Sheet No. 4.

page 280, and with the figures obtained by the use of the curves given in
Fig. 312.

The first step is to plot the magnetization curve. The figures for 6200, 7200
and 8600 volts are given in the calculation sheet, and the magnetization curve E
is shown plotted in Fig. 349a. Fig. 314 will serve for our graphic construction, but
is not to scale, as we will apply new values to the various vectors.

The armature ampere-turns per pole, 1^, are 5500. The terminal voltage E^
is 6900. It:fl^a is 350, lafa is 70, making Eg 7200. We find the approximate position

364

DYNAMO-ELECTRIC MACHINERY

Dltzy^y^pfy i<,.^. Type.ff^TT A,C.CMn Vm IMTOR nQf^4Ml /ff.PoUa Elec. Spec. ..-^

KVA.^^i^(?.; PF.:.fi.: PhweJ : WoXXs^^QQ, AmpB per ttr^/O. ; Cjcles .^.<?.....; R.PU^PQ ,. r,^ ^mps ;

HP^ Amps p. WTA.21Q . Amps p. br. arm ? T«mp rise ^S^.C ReguUtion./g.^JS/T « '.Soverload ZS.'^.Z flTS

Customer.^ ;

Order No ; Quot No ; Perf. Spec/VPsJ ; Fly-wheel effect

Fnuat
Air

D« L X R p M

K.V.A.

^6xfO^

Arm. A.T. p. pole....j^!5^*C?<?

Max. Fid. ti.T. fJOlfU?..

Armature.

Stat.

Dia. Outs.
Dia. Ins

o
o

Gross Length

Air Vents ^JLL

Opening Min.
Velocity

Mean

Net Lengthufi^ x-89

Depth b. Slots

Section . '^70 „Vol.
Flux Density.

20S

/SO

SOm.peir se<:.
£L3

/9 2

LossldSfip. cu.£^^TotaI
Buried Cu.?g<?<ZTotal
Gap Area ^7^00. Wts
Vent AreaaaeSfi; Wts
Outs. kx^7SJiQa. Wts \MsOgP

Noof Segs
No of Slots
K.

ISO.

Mn.Circ.

Section Teeth .
Volume Teeth-
Flux Density.

26,000

3^.700 4-JAQO
2.9.SQ0

2^.QOO

311 A

225

Loss*//:5.p.cu C^Total

Weight of Iron-

to
o
o

3
TJ
C
O

Star.

Cond. p Slot

Total Conds .-

Size of Cond. IZ^-X
Amp. p. sq.

.Throw

Length in Slots_2^

Length outside /^^Sum
Total Length '7e
Wt of i,ooo_^<?^TotaI
Res. p. i.ooo'3i2f^otal

Watts ^.m 7nLL2

Surface p. /5,

Watts p. Sq

_2ae

1 7,50 O

3A0Q_

U^^QCL

9J700

57'SOO

/^O Slots

<J5^

|c 5(

<--

•31

I
I

Field

RotOP.

Dia.

i Total Air Gap

Gap Co-eff. IC.

Pole Pitch ^ Pole Arc
Kr

u^e

< >

JTTK

eoo

O'SS sq^cm. _
'375

, iOSO ,

~'32'6Jf0s,

1322
—S5

/odd

085

omEi

lOPolms

<— 75 — ^

(-22 -9t

Flux per Pole-
Leakage n.l

hxea.f<95 Flux density
Unbalanced Pull

No. of Scg
No.of Slots
Vents
K. _

Mn.Circ.
X =

/se-e

c-e

f'^f

^±

sS^Tem

/9-a

22-6 X

/7.44>0

7o^

Section

Weight of Iron.

I Shunts

AT. p Pole nl^ad S350

A.T.p. Polef.LoadL//<?^i? .

Surface \^QQQIl, !

Surface p. Watt_L ^^5j)0SS, fh49Ctuai

I* R - ^ Q550 i

LR. 5Q*5\

Amps. I /^6

No. of Tumsu-«__!._Z5i^
Mean 1. Turn

2'05rh,

Total Length \*^AP L

Resistance SScoki, '

-9J{o¥

'6x/'24' X 7^^0 X '796^4^0

Res. pen,

Size of Cond

Conds. per Slot-
Total

'29^

/9X4-

^S-Q kg<

Length

Wt. per i.oooJEL..

Total Wt

Watts per SqCfrJ.'Oaa
Star or Mesh

'73^7 c^.

(QQQ

Paths in parallel

Magnetization Curve.

Core

Stator Teeth
Rotor Teeth

Gap

Pole Body ..
Yoke

Section

Langth

3Z^0d €
/295 20 S
690 13

6;^^volts.

A-Lp***

/3.ob6 4-0
/4,'000 35

A.T.

B. lA.T.pn»

~85dor~3 z
62 VAoddi 35_

4-0

3760
820

'^5/52

7;?/^?. volts.

pnnjAT.

7450

a^po

f6,2dd

75
50

6726

EFFICIENCY.

Friction and W ..

Iron Loss

Field Loss

Arm. &c. PR

Brush Loss

11 load.

Output
Input

Efficiency ^-

full.

2000
2091
956

.L.U

a5.ac?voit8.

i^?oc_/e9_

^^P- f9^3Q0

A.T.p.<(N

900

60_
163

^A^Q^P^5pd4S€ji9 PJI

A.T.

5260

UAjZZTPl

Loss Cur.

/•5

Mag. Cur.
Perm. Stat. Slot
.. Rot. Slot X

Zig-zag __

zx 75 X /'5 =225=^
I 77 >( 2/0x4. x22S^3-3
End 2-5 X 7a X4200=^ 7-3

^tO

Comnnutator.

Dia.
Bars
Volts p. BaxL
Brs. p. Arm _
Size of Brs. .
Amps p. sq. _
Brush Loss _
Watts p. Sq. -

.Speed

Sx/«

Amps : Tot. /0'$x/0^

Imp. V +
Sh. cir. Cur.—

Starting Torque
Max. Torque ^
Max. H.P

Slip

Powtr Factor

t92 3SOkOi

iv/ti

ALTERNATING-CURRENT GENERATORS

365

of the centre line of the pole by the construction given in Fig. 314, and by trial
and error find /». equal to 3100 ampere-tums, and I^d equal to 4500 ampere-turns.
We must now refer to Fig. 313, and we find that for a ratio of pole arc to pole
pitch of 0-71 the coefiicient K^ is 0-4. And

0-4x3100 = 1240 effective cross-mag. a.t.

Dividing 1240 by 75-5 turns per pole, we get 16-4 amperes exciting current, which,
from Fig. 349a, we see would give us 1950 volts generated in the armature. Thus
Ec in Fig. 314 = 1950, and gives us the true position of the centre line of the pole.
We now find that E is about 6900 volts, which requires 6350 a.t. per pole on the
field. If we add the demagnetizing 4500 ampere-tums of the armature to the 6350,
we get 10,850 ampere-tums per pole required at full load. Now compare this result

10000

VoU \

I '

f o E ^AJ^

OAil/l / r-r "—^ r— -"*" "^

ovWf ,-^7 „— • "^

Z' 1 ^y^

/ J y^

wVu f / / '

t tl -^

— ' -■/— ■ ■ ■' ■

WUQ t i y^

j_ y _^

O/l/l/l / ^ — — —— <■ -1

ZOW -+ y/*^

-T ^"^

t:^

iz^^ ± J: ± ^

HOOAmp

0 50 100 ISO 200 Amp,

Fio. 849a. — ^Magnetization curve, J?, and short-circuit characteristic, {jr. of a 2500 k.y.a.

S-phase generator.

with that obtained by the construction given in Fig.. 305. To generate 7200 volts
we require 6728 effective ampere-turns per pole. Set off the vector I^r to represent
6728, and the vector Iza to represent 5500, as in Fig. 305. The sum is 11,000 ampere-
tums, so that the difference between the results obtained bv the two methods is
not of great importance. Now find the ratio of short-circuit ampere-tums to no-
load field ampere-tums. The demagnetizing ampere-tums are 5500, and those
required to overcome the armature reaction are 230, making 5730.

Th-atio 1^ = 0-9.

Taking now the abscissa 0-9 in Fig. 312, we find that for 0-8 power-factor load
we require 70 per cent, increase in the field current. This gives us :

6350 X 1 -7 = 10,800 a.t. per pole,

which again is not very far from the mark. We see therefore that any of these
methods gives a result sufficiently near the truth.

CHAPTER XV.

ALTERNATING-CURRENT TUEBO-GENERATORS.

Altbrnati NO -CURRENT turbo-generators difier from alow-apeed machines both in
the design of the stationary armature and in the conBtruction of the revolving field-
magnets. The special modes of clamping the windings on turbo -armatures have
been considered on pages 119 to 131 ; and the manner of ventilation has been con-
sidered on pages 205 to 217. We will consider here a few points relating to the
revolving field-magnet, before passing on to consider the design of some machines
in detail.

Fio. S50.— Field magnet o( 8000 K.w. turbo-gmeralor, hsTing uUent pol«; speed 1800
a.F.H. (Wratlagbouse Co.).

The high speed of turbine-driven generators gives rise to very great centri-
fugal forces, which necessitate very strong construction in order to secure the
parts of the rotor. The number of poles on a high-speed machine of normal frequency
is necessarily low, and this leads to the use of wide and bulky field-coils, the support-
ing bf which makes the problem especially dif&cult. On the early turbo-generators,
salient poles, each with a single field-coii, were employed ; but it soon became
evident that the field-coil should be split up into small sections, each of which could
be independently supported in a more mechanical manner than was possible when
large f^gregations of insulation and copper were employed. Fig. 350 illustrates
a successful form of salient pole machine built by the Westinghouse Company
of America. It consists of two steel castings, extensions of which are forged down
to form the shaft. The main boilies of the castings are spigotted as shown, and
twelve bolts near the periphery hold the castings together. Four slots are then

ALTERNATING-CURRENT TURBO-GENERATORS 367

planed around each pole to receive the field-windings, whicli are wound directly
in the slots and insulated with mica. These coils are retained by means of trass
wedges closing the mouths of the slots, so that the whole field-magnet is enclosed
in metal. The method of ventilation ia clearly seen in the figure. This construc-
tion is exceedingly good from the mechanical point of view. It will be seen, however,
that the cross-section of the magnetic circuit is somewhat more restricted than it
would be in a cylindrical field-magnet (see Fig. 351).

From a theoretical point of view, leaving out of account all the difficulties
of supporting the parts mechanically, the ideal arrangement of copper and iron on
a fouF-pole field-magnet would be one which gives the greatest possible cross-
section to the iron paths, while the whole of the space between the poles ia occupied

Fia. 351. — Cyllndrlrol fteld-mtigiwt built up of punchlngi.

by the copper of the field winding. It will be seen that a cyUndrical field-magnet
with copper placed in slots near the periphery, more nearly approaches the ideal
arrangement than the salient-pole rotor.

There are several advantages to be obtained by placing the exciting winding
in radial slots. (I) Each section of the winding is well supported by the teeth.
(2) The cooling is good, because no part of the copper is very far removed from the
iron teeth, and the total coil surface is great compared with its volume. (3) The
efFective width of the pole is not merely the width of the smallest coil, but extends
across the pole pitch. Thus, while the copper is divided into sections and can
be worked at a high-current density, the space between the coils is not wasted, but
is used for the magnetic circuit. (4) It is desirable in many cases that the iron of
the magnetic circuit near the periphery of the field-magnet should be saturated.
The cutting away of the iron to make room for the copper is in this case a gain rather
than a loss. This will be seen more clearly when we consider the shape of the field
form of a cylindrical field-magnet. (5) The magnetic field-form can be made approxi-
mately sinusoidal, and results in a wave-form of electromotive force,* very near
to the true sine wave, both at no load and at full load.
•See ■' The NoD-Mlient Pole Tarbo-Altemalor," S. P. Smith, Jount. I.E.B., vol. VI, p. 662.

368

DYNAMO-ELECTRIC MACHINERY

The satiiration which occurs at the root of a salient pole is not as effective in
improving the regulating quality of a fiejd-magnet as saturation occurring near
the face of the pole. Where saturation occurs at the root of a pole, it will be found
that the ampere-tumB are very much increased on loads of low-power factor ;
because not only have ampere-turns to be added to overcome the normal satura-
tion of the pole, but extra ampere-turns must be added to overcome the excessive
saturation created by the leakage flux. Fig. 352 shows the form of the no-load
magnetization curve of a salient-pole machine, which had 20 per cent, of its field
ampere-turns expended on the iron at no load, 9000 volts. When full load
(cos</)=0-8) was thrown on the machine, the ampere-turns had to be increased
to more than double their value at no load, because the full-load magnetization

10,000
9,000
8,000
7,000
Z 6.000
6,000
4,000
3,000
2,000
1,000

—

-uS

r
1

10,000
9,000

8,000
7,000

•3 6,000

5,0QO
4,000
3,000
2,000
1,000

^^ '

0 20 40 60 80 100 120 140 160

EXOITINO CUKRENT.

FlQ. 352. — No-load and full-load characterlgtlcs
of salient-pole generator.

0 20 40 60 80 100 120 140 160

Exciting Currbnt.

FlO. 353. — No-load and full-load characteristics
of cylindrical field-magnet.

characteristic curved over so as to become almost horizontal at full voltage. Fig.
363 shows the general character of the magnetization curves on no load and full
load of a generator with a cylindrical field-magnet. The saturation occurring on
the surface of the pole has been adjusted so as to give 20 per cent, of the ampere-
turns expended on the iron at no load, 9000 volts. With this construction it would
be possible to obtain 9000 volts full load (cos<^ = 0-8) with an increase in the
ampere-turns of not more than 70 per cent.

The body of the rotor. There are three general methods of constructing the
body of the rotating field-magnet. (1) It may consist of punchings or plates built
upon a central shaft, as shown in Figs. 351, 220. (2) It may consist mainly of the
shaft itself, whose diameter is sufiiciently great to allow dovetail slots to be cut
in it, as shown in Figs. 354, 355, into which slots iron teeth are fitted. (3) The
whole rotor may be cut out of a solid cylinder of steel, as shown in Fig. 362.

Rotor built of punchings. The advantage of building up the rotor of steel
punchings or plates is that it enables the manufacturer to use rolled materials

ALTERNATING-CURRENT TURBO-GENERATORS 369

of great strength ; and the punching of slots and ventilating holes of the required
shape is a comparatively cheap process. The disadvantage is, that the diameter
of the shaft cannot in general be made great enough to give to the whole rotor a

PlO. 3H.— Iron p«rte of two-pola tnrbo ftaW-nU^net by A.E.O.

stiffness which will make the critical speed higher than the running speed. Most
tuibo field-magnets built up of punchings or plates have a critical speed lower
than the running speed. Very many successful machines have been made in this
way, and no difficulties are experienced in the balancing or nmning where the
proper precautions have been taken. Rotors of this type are illnstrated in
Figs. 220. 351 and 367.

Botor with dovetajl teeth. The second method of building up the rotor is very
well shown in the two-pole field-magnet built by the Allgemeine Elektiicitats

f two-pole tuibo fisId-nuwDBt b; A.E.O.

Gesellschaft, illustrated in Figs. 351 to 357, designed to run at 3000 revs, per
minute. Here we have a shaft of diameter sufficiently great to give the whole
rotor a critical speed higher than the running speed. In this shaft are milled dove-
tail grooves, as shown in Figs. 351 and 356, and into these grooves are driven

370 DYNAMO-ELECTRIC MACHINERY

teeth, BO as to secure the field-coils against the great centrifugal forces. With this
tTpe of construction, it is possible to build up the field-coils and teeth as a complete
whole, and to push them longitudinaUy ii)to place on the shaft ; oi, with a alight

PlO. 350. — Tvi>pa1e tnibo Rdd-nusnet by A.E.O,. thawing tba eietUng ooUi In poriUon.

modification of the construction, the field-coils can be put on one by one, beginning
with the largest coil, and the teeth inserted afterwaids. One great advantage of
this construction is that it enables each field-coil to be completely formed and

Fia. 3&T.— Two-pole tnibo Ocld-nugnct tor SOOO K.W. runnlog tl SOOO b.p.k.

insulated before it is put on to the rotor. Fig. 366 shows the field coils in position
before banding. It will be seen that the dovetails at the end of the rotoi form
ventilating ducts which supply air to the end windings and the ventilating holes
punched in the body of the teeth (see
Fig. 364). The ends of the windings are
secured by steel wire wound over a metal
casing, the whole rotor being finished in
the manner shown in Fig. 367. Drawings
of stator and rotor of a completed machine
are given in Figs. 376 and 377 (page 406).
Solid lotoi. Many manufacturers
prefer to construct the rotor of one solid
steel forging, and to plane out the slots
for the reception of the winding. In this
case it is convenient to provide ventilating
ducts immediately below the slots. The
form of the slots and ducts may be as
indicated in Fig. 368. This construction
l&.'Sl£^Su"otSSStid°' gi^es great lateral atiffness, and enable*
rotors of great length to be built, which
have a critical speed higher than their running speed.

One of the main difficulties in the design of cylindrical field-magnets lies in the
supporting of tbe field coils where they project at the ends of the rotor.

ALTERNATING-CUBRENT TURBO-GENERATORS 371

Botor windincB: Two-pole windings. Where the slota are radial and die
teeth are immovable, the only way of inserting the fieid-eoilB is by puttii^ them
in tnm by turn. The shape of field-coils naed on solid rotors with radial teeth is
shown in Fig. 369. In this case there are eleven coils per pale, and it will be seen
that while the cooling surface of the parts of the coils lying in the slots is exceedingly
great, the cooling conditions of the projecting ends of the coils where they are cloeely
huddled together aie not very good. The method of calculating the temperature

Fie. 35B. — Field colli of two-pole turbo fleld-nucnet.

rise on a winding of this t}rpe is given on page 227. A winding consisting of five
double coils is shown in developed plan and sectional elevation in Fig. 34tO. Com-
plete rotors of this t}rpe are shown in Figs. 221, 361 and 378.

Some makers have constructed very successfiil two-pole field-magnets with
barrel end-connectors (see page 116). A rotor of this type is illustrated in Fig. 220.
This constmction has a great deal to recommend it from the mechanical point of
view ; but on two-pole machines the end-connectors project from the active iron
very much further than with the " coil " type winding illustrated in Figs. 369
and 360.

Another type of two-pole winding which has been successfully developed by
the Westinghouse Company of America is shown in Fig. 362. Here the rotoi
cannsts of a solid steel for^i^ of cylindrical shape, in which parallel slots have been
cut in the sides and end. The winding consists of copper strap wound directly
in the slot, insulated with mica, and secored by means of bronze wedges. After

372 DYNAMO-ELECTRIC MACHINERY

this part is wound, flanges of bronze which cany the shaft are fastened at either
end of the field cylinder by massive screws. In the case shown in Fig. 362, the

shaft ends in a boss which has teeth machined in it not unlike a iaige bevelled
wheel. The bronze is cast around this boss, filhng the dovetails between the teeth.

ALTERNATING-CURRENT TURBO-GENERATORS

373

and making a good rigid connection. The rigidity of this construction is shown by
the fact that the critical speed is higher than the running speed, even when the latter
is as high as 3600 revs, per minute. Oeneratots running at this speed are built
of capacities as high as 5000 s.v.A.

a. 381. — Two- pole Of

it by Slemena Scbuckett Co.

BotOT windings: Four-pole windings. When there are four poles, the end-
connections of the windings are much shorter than on two-pole windings. There
is consequently much less likelihood of overheating. End-connectors of the barrel
form illustrated in Eig. 133 are quite suitable, and do not project too far from the
iron when the pole pitch is only one-fourth of the circumference. Fig. 363 shows
a finished four-pole rotor of 3000 s.w. capacity, the end windings of which are secured
by steel bands.

A type of winding which is very suitable for four-pole machines is that illustrated
in Eige. 364, 365 and 371. The conductors lying in the slots consist of simple
bars, suitably insulated, with ends projecting in the manner shown in Fig. 364.
These bars are connected in series with one another by end-connectors mounted
between two steel cheeks, which grip the end-connectoiB — in the same way as the
bars of a commutator are gripped — by means of V-rings insulated with mica.
There are several advantages in this type of construction. The cross-section of
the end-connectorg can, if necessary, be made greater than the cross-section of

DYNAMO-ELECTRIC MACHINEEY

Fio. 3M. — Fonr-po]« tuibo flehl-DUMnwt, bu wound, with (md-connecton
betveen >[«el ch«eki (British WhUd^uh Compuiy).

ALTERNATING-CURRENT TURBO-GENERATORS 376

the bare. As there is plenty of room for the steel cheeks, these can be made very
mAssive, bo that a higher factor of eafety can be obtained than in those constructions
where the amount of steel in the end bell is limited by the space available for it.
Moreover, if it is desired to replace the conductors in any one slot, this can be
done without interfering with other parts of the windii^. Fig. 365 shows the manner
in which the end-connectoie are mounted during the process of mann&cture.
VIgB. 371, 372 give detail drawings of this type of winding.

The fleld-form of cylindricaJ fleld-magneta. Dr. Stanley P. Smith, in a
paper before the Institution of Electrical Engineers,* has very fully invest^ted
the field-form of the cylindrical lotor, and has shown that when the winding space
occupies from 0-6 to 0-9 of the pole pitch the effect of all harmonics higher than
the fifth can be neglected. As the winding factor (see page 306) for the fifth harmonic
is only 0-19, and as the third harmonic is completely neutralized on a star-con-
nected, three-phase machine, the resulting terminal pressure is very close indeed
to a true sine wave. The paper gives the values of the harmonics for different
winding widths, both with and without saturation, and clearly sets ont the analytical

• " Tha Non-Mlient Pole Turbo-Altonutor and its CharsoteriBtioi," Joun. I.E.E., vol. 47,
p. 562.

376

DYNAMO-ELECTRIC MACHINERY

I I I I I I

*Sg - fiq!9u»p xn|j

iuipuiM JOQOH ^ uo(Qne(Mq«iO

ALTERNATING-CURRENT TURBO-GENERATORS 377

method of arriving at the resulting e.m.f.* The effect of armature reaction is
also fully dealt with. With any given cylindrical field-magnet the field-form will
change somewhat, as the saturation of the iron is increased. At low saturations
the curve showing the flux-density in the gap follows very closely the magneto-
motive force curve ; but as the saturation is increased, the comers of the
magnetomotive force curve are rounded off in the manner shown in Kg. 14, page
19. The manner of plotting the field-form for different excitations will be clearly
imderstood from Fig. 366, which is taken from Dr. Smith's paper. The first step
is to plot an air-gap-and-tooth-saturation curve, as described on page 78, the
abscissa being ampere-turns, and the ordinate the flux-density in the gap. If,
then, curves showing the distribution of magnetomotive force along the rotor
face are drawn for various excitations in the manner shown in Fig. 366, vertical
lines can be run up from these, and where the lines cut the magnetization curve
horizontal lines can be projected which give the flux-density at the corresponding
points on the rotor periphery ; so that the curve of flux-density can be plotted with
ease. The figure shows, in dotted lines, the fundamental sine wave, and also the
third harmonic. By taking several field-forms in this way, and calculating the
voltage generated in the winding, we can plot the open-circuit characteristic of the
machine as shown in Fig. 366. An example is worked out in connection with a
15,000 K.VJ^. generator on page 395. The field-forms and magnetization curves
are given in Figs. 373 and 375.

The speciflcation of A.C. turbo-generators. The main provisions in the
specification will be the same as for slow-speed generators. Clauses are sometimes
added to ensure sound mechanical construction, and in view of the cost of the plant
and of the very great importance of continuity of service, the specification is
sometimes made more elaborate than for smaller machines.

The model specification given below contains more clauses than are really
necessary. A variety of clauses are given in case the circumstances should call
for them ; but we must remember that it is always desirable to keep the specification
as simple as possible, so that a manufacturer may not be hampered in supplying
his standard machinery.

* The reader is referred to pages 305 to 316 on the subject of e.m.f. wave-forms.

378 DYNAMO-ELECTRIC MACHINERY

SPECIFICATION NO. 5.

15,000 K.V.A. THEIEE-PHASE TURBO-GENERATOR.

Bxtentofwork. 51. This Specification provides for the supply, delivery on
site, erection, testing and setting to work in the power station
at of two Turbo- Alternators (together

with the steam turbines, condensers, air-filters and auxiliary
plant described in the specifications issued with this one and
bearing an even date).

Hating and 52. Each of the turbo-generators shall have the character-

General • • •• -■ ^

Characteristlcg. IstlCS SCt OUt DClOW :

Normal output 15,000 k.v.a. or 12,000 K.w.

Power factor of load 0-8.

Number of phases 3.

Normal volts 1 1 ,000.

Voltage variation 10,000 to 11,500.

Amperes per phase 790.

Frequency 50 cycles per second.

Speed 1500 revs, per minute.

Regulation 22 per cent, rise with full load

0-8 power factor thrown off.

Over load 25 per cent, for 4 hours and 50

per cent, for 15 minutes.

Exciting voltage. 200 volts.

Temperature rise after| 45° C. by thermometer,

60° C. by resistance.

6 hours full-load run

Temperature after 4'
hours 25 per cent,
over load

See clause 62.

Puncture test 23,000 volts alternating applied

for 1 minute between arma-
ture coils and frame.
1500 volts alternating applied
for 1 minute between field
coils and frame.

Plan of Site. 53. Plan No. 1 attached to this specification shows the

proposed general lay-out of the power station and the position
of the new turbo-generators.

ALTERNATING^CURRENT TURBO-GENERATORS 379

54. The proposed general arrangement of the power plant ^J^^^^^^
is shown in plan in the accompanying Figure 1, and in
elevation in Figure 2.

55. The Power station is connected to the Railway AcceMiwiity.
by means of a railway siding, and a crane capable of lifting

40 tons will lift weights directly from railway waggons to
the central floor of the station,

or,

56. The power station has a wharf on the banks of the river

. A crane capable of lifting 20 tons
will lift weights from barges to the floor of the station. The
contractor must make provision for the lifting and handling
of weights greater than 20 tons,

or,

57. The power station is half-mile from the nearest railway
siding. The contractor must make provision for the carriage
of all parts of the machinery to the site in question, and for
this purpose he is invited to inspect the site and its approaches,

or,

58. The approach to the power station is along an alley-
way, one point of which is not more than 11 feet wide. The
contractor must arrange the parts of the machinery so that
they can be brought on site through the existing approaches,
or if any cutting away of brickwork should be necessary,
this must be made good at the contractor's expense,

or,

59. The turbo-alternator will have to be transported from
the entrance of the station over existing machinery, to the
place where it is to be erected. On account of the small head-
room, it may be impossible to do this while the existing
machinery is nmning ; in that case the bringing in the parts
of the new machinery will have to be done between the
hours of 2 a.m. and 5.30 a.m., and the contractor must
make allowance in his tender for any additional expense
which this will cause.

60. There is an overhead travelling crane in the power use of crane,
station capable of lifting 30 tons, which may be used by the
contractor at his own risk, when the same is not required by

the purchaser or his agents. The contractor must make

380

DYNAMO-ELECTRIC MACHINERY

General

Purpoeesof

Plant.

Temperature
rise on
over load.

Wave-Form.

Type.

Balance.

Factor o/
Safety.

Bearings.

provision for the lifting of any weights that are beyond the
capacity of the crane.

61 . The present power station supplies 3-phase power at
a pressure of 1 1 ,000 volts to the town of , where
it is utilised for the driving of cotton-mills and other factories,
for traction purposes and for general lighting and domestic
use. The turbo-alternators covered by this specification
are intended to supplement the plant at present installed,
and must be suitable in every way for the purposes aforesaid.

62. After a four hours' run with a load of 1000 amperes
per phase at 11,500 volts, p.p. 0*8, the temperature rise as
ascertained by increase of resistance shall not be such as to
make the maximum temperature in any part exceed the
value specified by the International Electrotechoical Com-
mission as a permissible temperature, having regard to the
nature of the insulation employed.

63. The wave-form of the e.m.f. at all loads shall be
approximately a sine wave, and at no-load there shall not be
any harmonic having an amphtude greater than 1*5 per
cent, of the fundamental.

64. The turbo-generators shaD be of the horizontal type
with revolving field magnets. The contractor shall state
the way in which he proposes to withdraw the field magnet
for inspection or repair.

65. The revolving parts * shaD be balanced with extreme
accuracy, so that when running only the smallest possible
amount of vibration is communicated to the bearings.
Means shall be provided whereby the balancing weights
can be easily adjusted.

66. At the normal speed of 1500 revs, per minute, the
rotors shall have a calculated factor of safety in every part of
not less than five. The revolving parts shaD, before leaving
the contractor's works, be run at a speed of 1700 revs, per
minute, without showing any signs of movement of the
component parts relatively to one another.

67. Bearings shall be of the self-aligning type, and shall
be so arranged that the bottom half of the bearing may be

Critical Speed.

* In cases where the purchaser wishes to insist upon haying the critical speed
higher than the running speed he may add the following clause :

The revolving parts of the machines shall be so constructed that the critical
speed is not less than 1800 revs, per minute. The purchaser shall be entitled to call
for the calculations as to the critical speed, so that he or his agents may check
the same.

ALTERNATING-CURRENT TURBO-GENERATORS 381

removed without raising the shaft more than 01 inch.
Liners shall be provided to facilitate the alignment of the
bearings. All bearings shaD be interchangeable. Bearings
ahall not be water-cooled. They shaD be lubricated and cooled
by a supply of oil. The oil-supply shall be continuous and
under pressure. Oil-pumps of ample capacity shall be
suppHed, capable of maintaining a constant pressure of not
less than 6 lbs. per square inch at the bearings. After
passing through the bearings the oil shall be passed through
strainers into an oil reservoir, from which the oil-pump draws
its supply. The oil shall be forced through a thoroughly
efl&cient oil-cooler before being fed to the bearings. An
independently-driven oil-pump shall be provided with each
turbo-generator for supplying oil during the starting up :
this pump shall preferably be steam-driven. A lip shaD be
cast round the bedplates and bearing pedestals to intercept
stray oil. The shaft shall be provided with very efficient oil-
throwers, and the whole arrangement within the bearing
housings for ensuring against the escape of oil shall be so
efficient that after a six hours' run no oil can be detected
on the shaft or anywhere outside the housings.

68. The magnetic design of the rotating and stationary Eddy-currenta
parts shall be such that no eddy-current is generated in the "
journals or bearings, even when the bearings are uninsulated.

The bearings shall, however, be so constructed that they can,
if need shall arise, be completely insulated from the bedplate
and oil-supply ; and if there shall be any evidence of the
existence of eddy-currents, the insulation of the bearings and
all other work necessary to overcome the trouble shall be
carried out at the contractor's expense. The bearings shall
be provided with suitable arrangements so that their tempera-
ture can easily be determined.

69. The shaft shall be of forged steel having a tensile shart.
breaking strength of 38 tons per square inch and having an
elongation of 18 per cent, measured on a test-piece not less
than 3 in. in length and 0-6 in. in diameter. The shaft shall
have no sudden variations of diameter. The journals shall

be ground and highly polished.

70. The bedplate shall be of exceedingly stiff construction. Bedplate.
and shall be arranged so that either stator may be erected

on either bedplate.

382

DYNAMO-ELECTRIC MACHINERY

Ventilation.

NoiBe.

Cables.

Foundations.

Framework.

71. The generators shall be completely enclosed and shall
be ventilated either by means of a fan on the rotor or by
means of an independently driven fan which shall be suppUed
by the contractor, together with its motor and all necessary
auxiUary gear. The motor, if any, for driving the ventilating
fan shafi be of an approved type. Suitable telltale arrange-
ments shall be provided for warning the switch board atten-
dants in case of any accident to the ventilating arrangements.

72. The generators shall not give rise to any more noise
than is observable in machines of similar size and speed
built according to the best practice.

73. The main cables from the armature to the switchboard
will be provided by the purchaser under another contract.
The contractor shall supply suitable terminals for the arma-
ture and field connections and shall supply all necessary
cables between the alternator fields and the exciter. He
shall also supply any necessary wiring to ventilating motors
and other motors, if any, suppUed by him for the operation
of the plant. After erection the contractor shall examine
all connections from the switchboard to the apparatus
suppKed by him and satisfy himself that such connections
are properly made. He shall be responsible for switching-in
and paralleling the turbo-alternators with the bus-bars.

74. The purchaser will provide all buildings, foundations,
cable ducts and trenches, and floor-plates for the same.
The contractor shall supply to the purchaser within four
weeks of the closing of the contract proper drawings, tem-
plates and materials required to be built into the foundations,
so as to enable the purchaser to proceed with the building
of the foundations without delay. If through non-delivery
of proper drawings, templates or material aforesaid any
alterations or additions to the foundations shall become
necessary, the cost of the same shall be borne by the con-
tractor. All leveUinc of the turbo-alternator, bedding and
grouting on the foundation shall be done by the contractor.

or,

75. The contractor shall supply with each turbo-alternator
a steel frame built up of suitable girders of sufficient stif&iess
to carry the complete turbo-alternator set when placed on
the foundations suppKed by the purchaser in the positions
shown in Figs. I and 2. This frame shall be levelled and
grouted in by the contractor.

Here may follow clauses Nos. 5, 6, 8 (or its equivalent), 10, 11, 12,
13, 14, 15, 16, 17, 18, 19, 20, 23, 26, or such of them as are suitable.

ALTERNATING-CURRENT TURBO-GENERATORS 383

CALCULATION OF A 15,000 K.V.A. TURBO-GENEEATOR.
11,000 Volts, Three-phasb, 50 Cycles, 1500 r.p.m.

The calculation given here may seem to be unnecessarily long and complicated.
It has been thought desirable to give the reasons for the various stages in the process,
and these are sometimes rather lengthy. In actual practice not one quarter of the
figuring here shown would be gone through by the designer, because he would
make short-cuts based on his experience of previous machines. Nevertheless, the
ultimate reasons for the dimensions chosen depend upon some such arguments as
those given here.

The design sheet is given on page 387, and the dimensions of the various parts
can be scaled off from the drawings given in Figs. 367 to 372.

The method of using the design sheet is in most particulars the same as in the
case of the engine-driven generator given on page 316. The main difference arises
from the circumstance that the rotor in this case has a distributed winding wound
in slots. The machine being totally enclosed and supplied with air from an inde-
pendent blower, we can make more exact calculations of the air velocity in various
parts.

The first point to settle is the diameter of the rotor. For a machine with such
a great output we will make this as great as is consistent with maintaining a good
factor of safety. A peripheral speed of 18,000 feet per minute is not an excessive
speed for a large turbo-generator, so we will try a diameter of 46 inches or 117
cms. The size of the air-gap is fixed from the regulating characteristics, and it
will be known from previous machines, or from such considerations as appear later,
that it ought to be about 1^ inches, or say, 3-2 cms. This gives us an internal
diameter of stator of 48| inches, or say, 1234 cms. We may arrive at a preliminary
figure for the length by taking a likely JD^l coefficient. If we adopt the type of
construction given in Fig. 371, there is no difficulty in making a large four-pole
turbo-generator (of 22 per cent, regulation on a load of 0*8 power factor) with
a JD^l coefficient no greater than 20,000 inches, or 320,000 cms. This would give
us a provisional length of 85 inches. Another way of arriving at the length is to
fix upon the number of conductors. A machine of this kind can be worked at about
1000 ampere-wires per inch of perimeter, so we may have about 150,000 ampere-
wires. The ciurent per phase is 790, so that the conductors may be about 190
in number. A more convenient number is 180. We can then have 72 slots with
five conductors per slot and two paths in parallel. We might, of course, have 60
slots, with three conductors per slot, but this would involve the provision of a
conductor to carry 790 amperes, which would be so big that it would have to be
stranded, and that would result in a rather weak winding from a mechanical point
of view. The eictra cost of doubling the number of conductors is such a very small
percentage of the total cost of the machine, that it is generally worth while to put
two paths in parallel when the current per phase is very great. Another reason
for putting two paths in parallel is that it enables us to have 72 slofcs instead of
only 60. Sixty slots would give us 2370 amperes per slot, which, though not an
impossible number, is not as good practice as 1970 amperes per slot. We must
not, however, get the number of slots too great on a high- voltage machine, or the

•
•

B
O

PE4

ALTERNATING-CURRENT TURBO-GENERATORS

385

copper space factor will be very low. With 72 slots we have a slot pitch of 5-35
cms., which is rather small but sufficient. Now find the value of A^B on the assump-
tion of 180 conductors. We will see (p. 396) that the value of Ke for this t}^e
of field is about 04. The number of revolutions per second is 25, so we have

11,300 =04 X 25 X 180 x A^B x lO"® ;
A^B = 6-3xl08 C.G.s. lines.

HUt %

h
X

liiiillllliil

Fig. 868.— End view of 15,0CX) k.y.a. turbo-generator.

The length of the iron must be great enough to give us room in the rotor to carry
this total A^B, and at the same time provide sufficient copper to carry the requisite
ampere-turns on the field. Now the armature ampere-turns per pole are 16,800.

In order to secure good regulation, we should make the ampere-turns at no load
some 50 per cent, more than this, and at the same time highly saturate the teeth.
We ought, therefore, to have some 26,000 ampere-turns per pole at no load (see
page 387). If we work the copper at 3000 amperes per sq. in. and have slots about

W.M. 2 B

386 DYNAMO-ELECTRIC MACHINERY

4^ ins. deep, we will find that we cannot make the ratio of iron space to copper
space much greater than shown in Fig. 371. In this figure we have a tooth 1*5
cms. wide, and a slot with a mean width of 1-7 cms. We have chosen a parallel
tooth and taper slot because it is an easy matter to draw the copper strap so as to
make good use of the space in a taper slot, whereas a taper tooth is not so eco-
nomical in room. A taper tooth becomes too highly saturated at the base, while
the top is worked at too low a density. A taper slot, moreover, gives us most room
near the perimeter, just where it is most useful. We have in this rotor 104 slots,
88 being wound and 16 unwound. If we had made fewer slots, we should have
improved the copper space-factor, but, on the other hand, we should have had less
cooling surface. The proportions shown are not very far from the best theoretical
proportions in this respect, though no doubt the output of the rotor can still be
increased by deepening the slots and putting in more copper. With 88 wound slots
and six conductors per slot, we get 66 turns per pole. It would be quite practicable
with the same type of construction to make eight conductors per slot, and thus,
reduce the field current, but the space factor would not then be quite so good,
and the construction of the conductors would not be quite so robust. An exciting
current of 700 amperes is not excessive for so large a generator, and can be easily
dealt with if the collector rings and brushes are made ample and well designed.
There is some advantage in keeping down the exciting voltage and the number
of turns on the rotor, because then the voltage rise in the rotor at the instant of an
accidental short-circuit of the stator is not so great.

Having arrived at the size of the rotor teeth, we fix the amount of saturation
by considering how many ampere-turns we wish to expend on the iron. In order
to give the field distribution the form depicted in Fig. 14, we ought to expend
about 20 per cent, of the no-load ampere-fcums on the iron. Let us say 5200 ampere-
turns on the teeth, which have a length of 10-4 cms., giving us 500 ampere-turns
per cm. The apparent flux-density will depend upon the amount of slot and vent
space in parallel with the iron ; in other words, upon Kg. We may assume a Kg
of 2-5 in this first approximation, and from Fig. 47 we find that with 500 ampere-
turns per cm. we have an apparent density of 22,500 C.G.S. lines per sq. cm. Divid-
ing this into 6-3 x 10® we get about 27,000 sq. cms. for the area of all the teeth.
This gives us a net length of rotor iron of 173 cms., or, allowing for ventilating
ducts, say 200 cms. In actual practice the process of provisionally fixing the length
would be much shorter than given above. Having fixed the number of slots in the
rotor and their width, we would assume a density about 22,500, and arrive at once
very near a suitable length. The final adjustment of the saturation can be carried
out by changing the number of ventilating ducts, or inserting iron in the empty
slots. The length shown in the drawings is 204 cms., so we will proceed with the
calculation on that basis. If we have enough room for copper and iron on the
rotor, we always find in turbo-generators of good regulation that there is plenty of
room for copper and iron on the stator.

To get room for copper on the stator we have only to choose a slot of sufficient
depth. The increase of the self-induction of the armature with the increase in
the depth of the slot is, in fact, an advantage rather than a drawback, for the self-
induction of the armature of these big machines is generally lower than we wish

ALTERNATING-CURRENT TURBO-GENERATORS

387

Dau 6Mor^hi^l2 Type Tur^Q ...AClCBN •¥M MOTOW WQTAWV ....4. Pole. Elec Spec, n?

KVA/A:POO; P.F.:fi..; Phased : Vo\\sfflQQQ.tPM»fff^., Amps per ter..Z9.C>...; Cycles..^^. ...; R.PM./^.Qff..., Rotor Amp»

H-P: Amps p. coad. 3fii5 . Aipi p. hr, ■>- -^^./HPS-Temp nte .45.?.C. Regulation. ^<^%/?i^.*:iS. Overload J^S%.^Ao</.f:>

Customer PQtyj^.P^^ UQHJ QQ. . Order No.

Quot- No..

Perf Spec Fly-wheel effect

^;^C.rcum 3gg; Gap Area 7y<?(?^ ^^ ^ ^.^ ^ j^ ,^^^ A^Zl?C?^ ;Cl^m.3.7^.

D- L « R P Hi
K.V. A. ..

^3'/xi0^

Arm. A T. p. ^\^J.S.800. Max. Fid. hT.ieS,.000

Armature

Stat

Mean

Vol
Total

Dia Outs

Dia Ins ,

Gross Length ._

Air Vents ^P'

Opening Min ''P^ Mean
Air Velocit) _ -
Net Length j!6^
Depth b Slots. -
Section 47*^_

Flux Density

Loss-^?<L5p cu CHL
Buried Cu/iieP^ Total
Gap Area 72QQ0 Wis

. 2ja I

J23-4 \ _

2-5sg trt\

26^/0^

11,000

/2^ooq\
ni^66d\i66.ooo

l$7.^^

VentArea-?4^<2<W:\Vts (Z^uPOO

Outs. Arca/^ft^WWts

f^^QCfC^75I^OOO

No of Segs
No of Slots
K. _:?lJL^

Section Teeth _.
Volume Teeth _
riux Density

/^iMnCirc!
72\^2AX=\

445
175
270

. 36J900 I

564^000^
f6,200

Loss-^tf p cu C^. Total ^SJ)00\ __

Weight of Iron _... _

-^A^gj^

Star

Cond. p Slot

T.)tal Conds

Size of Cond./l^^- a
Ainp. p. sq.— ^^.

_ .Throw l^r^-^:|?^

8rC.

/•4

Len^h in Slots -^iP^f

Length outside ^MSurnL^" 44m

Total Ungth \.S00 X

\Vt. of \%oMyj> Total JL^PM-
Res p l?ooo-^i',5.Total ,_:^^2
Watts \y.jriSLre 98

Surface p
Watts p. Sq.

metre

0-4-

f9S0
^P5(_
05f

cms

?-? Slots

> i"X-

3!E

9-/

V 1/

*CZs5^

< >

msuts

8x3-2 X f '04- X 7600^2/000

Field

Rotor.

Dia.

\ Total Air Gap

Gap Co-efl. K^

Pole Pitch.S4^ Pole Arc

Flux per Pole.

//r

JS_i2,

/ 04

6€

(04-^/0?
H54^I0^

Leakage nMtP§t^ {XQ: 114
AreaflflflflLFlux density .J^MjSOO
Unbalanced PulL_ ;

No. of Scg.l_J I Mu.Circ. |— ^H

No.ofSlotsi_^4

Vents ■ij!^: Q^

x/-7=

/77

/56 r

K, ^S^^J^t\on254QQ^J4^Q_±,2j^QOO^
Weight of Iron fiun<:hings\.Ql^<ikg.^

Shunt. SttHa*. ' Com*n.

A.T.pPolen.Load -^^-^^ __

A.T. p. Pole f .Load .^7J>Q0 \

Surface

Surface p. Watt ,

I* R tO6j0OO

I. R. ' ZllPP. ".

Amps. 400n.L ■ 7/gy:/.

-^<

'50

3^1

XJ76Qm.:2LM

No. of Turn*
Mean 1. Turn
Total Length

Resistance

Res. per Tooa^^^\'ifSan<^i076 ■

It'Sji^cnuandlHZ^ sq. c/ft.

Cond,

Size of

Conds. per Slot— ,.

Total

Length

Wt. per 1,000.
ToUl Wt.
Watts per Sq.£.^

Star or Mesh

Paths in parallel

9m.

20QQ_

sfot

Magnetization Curve.

fO.OOO.yo\i&.

//,.Q.O.Q\/o\ts.

/ZOQO.^oMs.

Conn mutator.

Section , LCfigtH

A.T.P-nn

Core

Stator Teeth ^^ /O \fjpQ

Rr)tor Teeth A^_^v5^

Gap

Pole Body \ | 4.

Yoke _^_

"^•M^&AQQ

'50\ VPJSA

A.T.

200

/6^00 44>

A.T.p«i»

2ieP_ 22^Q0 490

7800' ?LOOp

AT.

:A.T.P«m A.T.

4-pg
^7ob

/Z70O[ 80 I 800

9SQ0_

Bar.

84dd

940

22.700

2/.430

msoo

crriciENCY

Friction and W —

Iron Loss

Field Loss

Arm &c. I-R

Brush Loss

IJ load.
160

tit

Output
Input

539

'"K*"- —7—

Efficiency y^

Full.
i60

171

120
3/'

502

/5,000/2.000
f5.63Sy2,502

965 96

/60_
17/

nq_
/6

J
160

'JLL
'96

- >

4—

4^79 455

9000 6000

9479 64J55

I6p

^-.
2

3000

3^M.

67-5

3SJ500

Dia

Bars _
Volts p
Brs. p. Arm
Size of Bis.
Amps p. sq.
Brush Loss
Watts p. Sq.

.Speed

Mag. Cur
Penu. Slat. Slot
Rot. Slot X
.. Zig-zag

Loss Cur.

2 X
177
End

^/S.

X
X X

Amps . Tot.

; X. =

; r- =

-f

Imp. V

tih. cir. Cur

Starting Torque
Max. Torque ._

Max. H.P

SUp

Power Factor

388 DYNAMO-ELECTRIC MACHINERY

it to be. In order that the armature current which will flow, if the machine is
accidentally short circuited at full voltage, may be kept within reasonable bounds,
it is well that the slot-leakage flux and end-leakage flux of the armature conductors
at full load should be equal to about 10 per cent, of the main working leakage
(see page 126). In large turbo-generators the main working flux per pole is so
great that unless some special provision is made for increasing the permeance of
the slot-leakage path, the armature stray field will be only a very small percentage
of the whole, .and the forces on the armature conductors at the instant of short
circuit may be excessive. It is therefore good practice to deliberately increase
the slot leakage. This can be done by making the slots of the shape shown in
Fig. 370. Incidentally we gain two points of advantage with this construction.
We have the armature coils well removed from the rotor, so that there is less fear
of a flash between rotor and stator. We provide a very useful cooling surface at
the head of every tooth, and allow more air to pass from the ends of the machine
to the middle than would be possible with ordinary slots. Observe that it is better
to have the mouth of the slot wide and the tooth head fairly long, than to have
a narrow mouth and a short tooth head, for though the leakage flux at full load
might be the same in both cases, the leakage at ten times full-load current will
be more, the wider we make the leakage path. We are, in fact, aiming at pro-
viding a leakage path which at ten times full-load current can carry, a flux equal
to the main working flux without undue saturation.

It is well to work out the leakage flux at full load. This can be done approxi-
mately from Figs. 369 and 370 as follows :

The flux passing from the head of one tooth to the head of the next per centi-
metre length of iron for 1 ampere total current in the slot is

5 1
l'25x^ o=3'5 c.G.s. lines.

1 -o

The effective flux passing across the slot under the same conditions is

In addition, we have some flux passing along the air-gap in a circumferential
direction ; this is equal to

l-25x^^„^^=0-77;
5-25

3-5 + 1-54 + 0-77 = 5-81.

The maximum value of the current in the slot at full load is

2i X 790 X 1 -41 = 2780 amperes.

The slot-leakage flux per pole is therefore

5-81 X 2780 X 204 X 2 = 6-6 xlO« C.G.s. lines.

As the working flux amounts to 104 xlO®, the slot leakage amounts to 6-35%.
Next, take the leakage aroimd the ends of the coils. This cannot be calculated
with any degree of accuracy. We may employ the formula given on page 426 for
the end leakage of induction motors,

Ia<i>e = Klx(Ip + Ov) X virtual A.T. per pole.

ALTERNATING-CURRENT TURBO-GENERATORS 389

The arrangement of the windings most closely resembles the case where we
have a concentric winding on the stator and a squirrel cage winding on the rotor,
so that Ki, from Table XVIIL, p. 427, is 2-8. Taking Jp = lll cms. and 0^=28 cms.,
we have for the full load ampere-turns, 11,850,

2-8 X (111 + 28) X 11,850=4-6 x 10» c.G.s. lines per pole.

Adding the slot leakage, we get

(6-6 + 4-6)10» = ll-2xl0«,

or approximately 11 % of the working flux.

The heads of the teeth are made wider than the body of the teeth by an amount
sufficient to give mechanical support to the coils. One advantage of wide heads
is that the iron loss is lower than if the heads were made narrow and long. A long
head, on the other hand, brings the armature slots on a larger diameter, and allows
a rather wider slot to be used than would be otherwise possible.

In fixing upon the size of slot, it muBt be remembered that plenty of room must
be allowed for insulation between turns. Although the normal running voltage
between successive turns is only 110, the insulation should be able to resist a puncture
test of 4000 volts. A good plan is to place a strip of micanite 1 mm. thick between
each turn, and in addition to this there will be two layers of half-lapped linen tape,
so we must add 1 -5 mm. to the depth of each conductor. The current per conductor
is 395 amperes. The size of conductor that we must employ will depend upon the
cooling conditions. Here we have 11,000 volt insulation and a high current per
slot, so it will be found that we cannot work at a high-current density. To find the
permissible current density, we must make a rough guess at the cooling conditions.
We know that the cooling surface of the coil in the slot will be about 2000 sq. cms.
per metre length. Suppose that we allow 18° C. temperature difference between
the inside and the outside of the coil. With the teeth at 50° C. that would mean
a temperature of 68° C. for the copper. The insulation will be about 0-4 cm.
thick, and the heat conductivity 0 0012 watt per sq. cm. per degree per cm.

13 • -1.1 .. 00012x18 ^^-,
rermissible watts per sq. cm. = ^r^j = 0'054.

This allows us 100 watts per metre length of coil, and as we have 5 conductors,
each canying 395 amperes, it is easy to calculate that the resistance per metre
when hot should not be more than 0 0001 28 ohm. If we choose a conductor with
a cross-section of 1-65 sq. cms. it will be about right. This has a resistance of
0105 ohm per 1000 metres at 20° C, so allowing for 50° C. rise we have the
watts per metre length of coil

395 X 395 X 0000105 x 1 -2 x 5 = 98 watts.

The actual mean perimeter of the insulation works out at 19*5 cms., so that
the cooling surface is 1950 sq. cms. per metre, giving the required sq. cm. per watt.

Having obtained our cross-section, the next point is to fix on the external
dimensions. We would, in practice, be guided in this by considering what slot
dies we had available, but in the absence of any such consideration we will make
the width of the strap as great as we can, so that the depth may be as small as
X>ossible. In this case we cannot make the width greater than 1 -4 cms. or we will

390 DYNAMO-ELECTRIC MACHINERY

make the teeth too narrow. So our conductor, if in one piece, would be 1 4 by 1 -28.
Now, we see from Fig. 167 that if we used a solid conductor as deep as 1-28
cms. near the mouth of the slot, the eddy-current loss on it would be very excessive.
We have a =0-78 and /= 1-28, so that a/'=l. Now, from the curve m=5, we find
that the loss in the top conductor will be 7-5 times as great as it should be. We
have therefore made the top conductor and the one next to it of stranded copper,
and the rest of the conductors we have divided into two parts, slightly insulated
from each other. All the end connectors we have made of double straps, twisted
on themselves midway along their length, as shown in Fig. 370. There are two
objects in twisting the end connectors. In the first place, they interconnect the
top and bottom halves of conductors l3dng in difEerent slots, and so neutralize the
eddy current which would otherwise circulate in these two halves. Secondly,
by the twist we neutralize to a great extent the eddy current which would be
generated in each end connector itself. The total maximum current in all the
end connectors amounts to 29,000 amperes. This will set up a very strong field
in the body of the copper connectors, and it is desirable that they should be
laminated as much as possible. Stranded copper connectors would be better
electrically than the straps shown in Fig. 370, but they would be rather weak
mechanically.

Thus, we arrive at the arrangement of conductors shown in Figs. 369 and 370.
Taking into account the requisite insulation (see page 201) between turns and the
outside insulation, we arrive at the dimensions of the slot shown. It will be seen
that, in addition to the retaining wedges at the mouth of the slots proper, there are
wedges bridging across between the heads of the teeth. These are to prevent
excessive noise and churning of the air. It is best to make these wedges in short
pieces, each no longer than the thickness of a packet of iron punchings, so that the
air has easy access to the cooling surface afiorded by the heads of the teeth. The
width of the conductor has been chosen so that sufficient iron is left in the teeth.
This cannot be finally checked until the number and size of the ventilating ducts
is fixed.

A good rough rule for settling on the number of ventilating ducts on big turbo-
generators is to allow one duct for every 1 J inches of iron on a 50-cycle generator,
and one duct for every 2 inches of iron on a 25-cycle generator (see page 253). The
size of the ducts will depend on the amount of air that must be put through the
machine, and the pressure available. Where it is intended to employ an indepen-
dent blower to give the air supply, fairly narrow ducts can be used, as it is a very
simple matter to increase the air pressure, if it is found that too little air is passing.
A handy formula for calculating the amoimt of air required is the following :

^ , . ^ . . , 0-85 K.W. loss

Cubic metres of air per second = -. -. — ; rrr-

^ temperature rise of air, L>.

Or, in other units,

^ V . r . •. 1*78 X 103 XK.W. losg

Cubic feet per minute = . ^ — -. ttt'

^ temperature rise of air, C.

In these formulae the volume of the air is supposed to be measured at 20** C.
At 60° C. the volume will be U % greater.

ALTERNATING-CURRENT TURBO-GENERATORS 391

We know from previous experience that the losses in the 15,000 K.V.A. generator
will be about 500 K.w. If we allow an average temperature rise of 25° C. for all
the air going through, we have

0-85x500 ,^ , . , ,
p,_ - =17 cubic metres per second.

25 ^

On the calculation sheet it will be seen that we have allowed 16-5 cubic metres
per second. This is equal to 35,000 cubic feet per minute.

Velocities of air in various parts. If, now, we take 50 ventilating ducts, each
0-8 cm. wide, we will have, half-way between the internal and external diameters
of the stator, a total area of path of 2-5 sq. metres. With an air supply of 16-5
metres per second, we will have a mean velocity of 6-6 metres per second. This
is quite a suitable velocity. The minimum opening of the ventilating ducts (that is,
near the slots) is 1 03 sq. metres, giving a velocity of 16 metres per second. This
is fairly high, but not excessive. The total area available for the passage of air
in an a2dal direction, along the air-gap and along the rotor ducts and empty slots,
is about 0*5 sq. metre, so that the velocity of the air entering the air-gap at each
end will be about 30 metres per second, and the velocity along the rotor ducts
will be rather higher than this.

Having settled on 50 ventilating ducts, each 0*84 cm. wide, we get by sub-
tracting 42 from 204, 162 cms. of punchings and paper. Multiplying by the factor
0-89, we get 144 cms. of solid iron.

Flux-density in the teeth. We are now in a position to check the cross-section
of all the teeth. As explained on page 322, we find the maximum density in the
teeth by merely dividing the total AgB by the cross-section of all the teeth. Imagine
a circle drawn through all the teeth, which has a diameter of 142 cms. It has a
circumference of 445 cms. This we call on the calculation sheet Mn. circ, or mean
circumference. From this must be subtracted the sum of the widths of all the
slots, or 72 X 2 -42 = 175. This leaves us 270 cms. of iron all the way round. Multiply
this by 144, and we get the total mean section of the teeth as 38,900 sq. cms. Divide
6-3 X 10® by this, and we get 16,200 for the mean flux-density on the teeth. This
is rather high for a big generator, but not too high. The volume of the teeth is
obtained by multiplying the section by the length. The loss per cu. cm. can be
obtained by referring to Fig. 29. For very special machines, however, we can,
as explained on page 52, by extra care make the iron loss considerably less than
that given by the curves on Fig. 29. In this case we take the loss on the teeth
at 0-08 watt per cu. cm., which gives us a total loss on the teeth of 45 K.w.

Depth below slots. This dimension depends upon the total flux per pole.
Dividing 6-3 x 10® by 4, and multiplying by the form factor Z/=0-66, we get the
working flux per pole 104 x 10^ c.G.s. lines. If we allow a flax-density of 11,000
per sq. cm., we shall require in each half of the path for this flux 4740 sq. cms.
Dividing 4740 by 144, we get 33 cms. for the required depth of iron.

This depth is often fixed in practice by the bore of some existing frame, and
sometimes the flux-density will be higher or lower than 11,000 to fit existing parts.
Where no such restriction exists, one employs a density of 11,000 for 50-cycle
generators and 12,000 for 25-cycle generators. The volume of the iron behind the
slots is obtained by multiplying the area 4740 by the mean circumference 585 cms.

392 DYNAMO-ELECTRIC MACHINERY

This gives us 2-8 x 10* cu. cms. We may take the losa at 0-045 watt per cu. cm.,
so that the loss behind the slots is 126 K.\v. Adding this to the tooth loss, we get
171 K.w. The total length of armature coils buried in the iron is
72x2 04 = 147 metres.

FlO. SflB.— S«ctk>Ti of tnd wladlns ot IS.OOO K.v.i. turbo-gtmnitor, tiiowiiig the clsmp
bolted to tha eitenial tIdc o( fenaer.

Multiply this by 98 watts,, and we get 14-5 K.w. tor the buried copper losses.
The total losses to be dissipated by the surface of the stator is therefore 185 K.w.
Coolinc of tha stator. We have now to consider how we will get rid of all the
heat generated by this lost power.

First, take the inside cylindrical surface or gap-area. This is 77,000 sq. cms.

1-hMi
333

394 DYNAMO-ELECTRIC MACHINERY

It is shown below that the mean-temperature rise of the air in the air-gap will
be about 16^ C. Taking the iron at 40° rise, we have a difference of 24° C.

17 = 92 metres per second.

Therefore watts per sq. cm. =0-74,

77,000 sq. cm. X 0-74 = 57,000 watts.

Next, take the cooling from the walls of the ventilating ducts. The total area,
counting both sides of the ducts and allowing for spacers, is about 2,400,000 sq. cms.
The mean velocity in the ducts is 6-6 metres per second.

A,=00007x 6-6=00046.

Before we can estimate the watts per sq. cm. dissipated by the surfaces of the
ventilating ducts, we must find the mean-temperature rise of the air in the ducts.
We must first ask what amount of heat is received by the air before it enters the
ducts. We have yj^j^ pj^ loss = 106 k.w. (see p. 387).

Windage of rotor = 125

Heat from inside stator= 57

Armature connectors = 17

= 305 K.w.
0-85 X 305

16-5 "^^^ ^•
mean rise of temperature of air before entering the ventilating ducts. Now,
as we only have between 100 and 120 K.w. to get rid of from the ventilating ducts,
this will make a further rise of

0-85x120 ^ori
16-5 ^^ ^-

The mean temperature rise of the air after it has passed half-way through the
ducts is 19° C. above the outside temperature. If, now, we take the temperature of
the surface of the ducts at 35° rise, we have a difference of 16° between air and iron.

Watts per sq. cm. = 16 ; A,, = 16 x 0 0046 =0-074,

0 074x2-4xlO« = l 75,000 watts.

We have therefore very much more cooling surface than is necessary to carry
away the heat generated from the calculated losses. In fact, the mean temperature
rise to be expected is only 19° -i- 6° = 25° C. We must, however, remember that
the temperature in the middle of the machine may be some 5° or even 10° hotter
than at the ends, and the iron loss may be somewhat higher than we have calculated ;
therefore the margin which we have allowed is desirable.

A certain fraction of the heat generated is dissipated by the end plates and
conducted into the frame, whence it passes to the air circulating through the frame.
For large turbo-generators, we may allow about 0-1 watt per sq. cm.* of external
surface for the heat dissipated in this way. The external surface is 190,000 sq. cms.,
so we have about 19 K.w. dissipated by conduction into the iron frame and to the
end plates.

* The reason for allowing a smaller rate of cooling on the external surface than on slow-
Mp<^ed machines is that the ratio between the surfaces through which the heat is conducted and
the wliole external surface is less than on slow-speed machines.

ALTERNATING-CURRENT TURBO-GENERATORS

395

The various quantities of heat dissipated from the gap-area, the vent-area
and the outside area for a temperature rise of 40° C, are entered in their
respective places on the calculation form (page 387). The sum is 251 K.w, As the
calculated loss to be dissipated from the iron parts of the stator is 186 K.W., the
temperature rise to be expected is lower than 40° C.

The other figures for the armature given on the calculation sheet explain them-
selves.

The length of the air-gap is fixed so as to give us about 21,000 a.t. per pole,
that is to say, some 25 % greater than the armature ampere-tums per pole. The
maximum flux-density in the gap is obtained by dividing the gap area 79,000*
into 6-45x10®. This gives us 8170 C.G.s. lines per sq. cm. We then have

0-796 X 3-2 X 1-04x8170 = 21,600 ampere-tums on the gap at 11,600 volts.

As the air-gap is so great in comparison with the opening of the slots and venti-
lating ducts, the gap coefficient is nearly unity. The working flux per pole is
obtained by dividing 6-3x10® by 4 and multiplying by 0-66 the form factor Kj,
In making out the tables for the magnetization curve, it should be remembered
that the maximum density in the gap is not quite proportional to the voltage,
because the coefficient Ke changes slightly as the saturation increases. In practice
we change the constant by a small amount, which can be judged from experience.
If we wanted to be very accurate, we would have to make a plot of the field-form
at two or three voltages, as shown in Fig. 373, and determine K^,

In taking the cross-section of the teeth, we must not forget the area of the spacers
in the ventilating plates. These have a total section of 1400 sq. cms., making total
section of iron 26,800 sq. cms. If we multiply the mean circumference 333 by the
total length 204, we get the total section of air and iron, and dividing this by 26,800
we get j&, = 2-54. The high value of Kg makes the apparent flux-density in the
teeth very much higher than the actual, and has a great influence on the number of
ampere-tums required for the teeth, as can be seen from Fig. 47. If we work out
the ampere-tums per pole for various flux-densities in the air-gap, we arrive at
the figures given below :

B in air.gap.

Apparent B In
teeth.

Amperctums
per cm., irt= 2 '54.

Ampero-tnms
on teeth.

6,000

17,200

625

7,000

20,500

200

2,080

8,000

23,000

600

6,250

9,000

25,800

1,360

14,050

10.000

28,750

2.300

24,000

For B = 8000 in the gap, we have

0-796 X 3-2 X 1-04 x 8000 = 21,050 ampere-turns on gap.

This gives us the position of the air-gap line shown dotted in Fig. 373. From
this line we set ofl the ampere-turns on the teeth as described on page 78, and
obtain the " air-gap and tooth " saturation curve. In order to find the field-form

* The gap-area for this purpose is taken at 79,000 to allow for the fringing at the ends of the
rotor.

396

DYNAMO-ELECTRIC MACHINERY

^ e( CO N

•*■ oico^'o«oN«)Oig5<52

ALTERNATING-CURRENT TURBO-GENERATORS 397

9 excitations, we eet oS the trapeziume which give the distribution of
magnetomotive force (see page 375). These are shown at the base of Fig. 373
for 22,000, 26,500, 33,000 and 47,000 a.t. respectively. Running up the ordJnates
for the ampere-tums on each tooth until we strike the " air-gap and tooth " curve,
and then along horizontally as shown in Figs. 366 and 373, we can plot the field-
forms shown. By means of a planimeter we at once ficd K/, and Kf can be found
by the method described on page 28. Another way of arriving at Ke is to take the
value of the voltage coefficients as determined by Dr. S. P. Smith,* and find its
ratio K^ to the voltage coefficient for a sine-wave field-form. Now £, for a sine-
wave field-form is 0-39 (see page 25), so that 0-39 xX^r =A',.

It will be found that, for field-forms of the general shape of those shown in Fig,
373, there is a fairly close relation between the value of Kt and Ky, so that after
we have worked out a number of cases we can plot a curve giving the relation as

>i A«lf1-(onnB of the lenenl Khap«

shown in Fig. 374. It is then only necessary to find Kj by means of a planimeter
(see page 16), and read ofi the value of K, from Fig. 374. The change in the values
of Kf and K, as the excitation is changed will be seen from the curves plotted at
the right-hand aide of Fig. 373. Knowing the maximum values of B for various
excitations, and the values of Kf, we can now find the voltage by means of the
formula Volts = K, x fip, x No. of conductors x ^^ x B.

For instance, at 22,000 volts, we have

VoIt8=0-382-x 25 X 180 X 79,000 x 7200=9800.

We can now plot the no-load magnetization curve as shown in Fig. 376.

The amount of iron in the rotor teeth and the length of the rotor teeth are
adjusted so as to absorb about 20 %^ of the ampere-tums per pole at no load. In
this machine 5100 ampeic-turns are absorbed on the teeth at 11,000 volts. This
amount is so great that the ampere-tums absorbed by the armature teeth and

398

DYNAMO-ELECTRIC MACfflNERY

core can in general be neglected. We, however, give the figures for the armature
teeth on the calculation sheet, though the possible error in the figures for the rotor
teeth make these small figures of little value.

i&poo
HWO

13J0O0
tItflOO
11,000

taooo

S(P00
8,000

tooo
aooo

5J0OO
4^000
5.000
StfiOO
ffiOO

'N'

1 1 1 1 1 1 «

)Lts and Flux -density

b-/

^ * •

^**

P .''

■A

/ 1

J 1

t
t

/ /
/ *
/ /

/ /

*
t

/ 1

m.000 20000 jaooo ^oiooo
Ampere 'Turns perfhle

SQOOO

Fia. 876. — No-load and full-load magnetisation curves of 15,000 K.Y.A. generator.

The field leakage depends mainly upon the permeance of the field slots. Apply-
ing the ordinary rules, we find that 1 ampere passing in a slot creates about
5-5 C.G.S. lines per cm. length. At 400 amperes we have

400 X 6 X 5 -5 X 204 X 2 = 5 • 4 X 1 0« .

To this should be added about 1 x 10^ c.G.s. lines for end leakage.

It will be seen from Fig. 371 that some wedge-shaped pieces of iron have been
inserted in the slots at the sides of each pole. These iron wedges are so proportioned
that they will carry the no-load leakage when saturated to the same extent as
the teeth are saturated by the working flux. There is therefore no increased satura-
tion due to leakage at normal-voltage no-load. At full load, however, the leakage
is increased to 114x10' c.o.s. lines per pole; the difference 5x10* so highly
saturates the teeth in the centre of the pole that there would be required an increase
of 4000 in the ampere-turns on the teeth from this reason alone, were it not for
the change in the value of Ke. By plotting the field-form under the new con-

ALTERNATING-CURRENT TURBO-GENERATORS 399

•

ditions by a process of trial and error, it will be found that, with the ampere-turns
increased to 47,000 per pole, the value of Ke goes up to 0-425, and this reduces the
extra ampere-turns required for the centre teeth to 2500. By taking two or three
points on the saturation curve, and investigating in this way the effect of the
increased saturation, we get the dotted curve NN' for the magnetization curve
with increased saturation on load.

Having obtained this curve NN\ the plotting of the full-load magnetization
curve is carried out exactly as described on page 386, and is given in Fig. 375. We
find that with an inductive drop in the armature of 10 % (see p. 389) it is necessary
to generate 11,700 volts in order to get 11,000 volts at the terminals at full
load, 0-8 power factor. Taking the ampere-turns required for 11,700 volts
from the curve NN\ and compounding these as in Kg. 305 with the 15,500 ampere-
turns of the armature, we arrive at 44,500 ampere-turns per pole at full load,
0-8 power factor. It is well to allow some margin on this to allow for the iron being
more highly saturated, as would be the case if the punchings were not very tightly
packed. We have taken 47,000. This gives us an exciting current at full load of
710 amperes.

The calculation of the cooling of the copper in the rotor slots is straightforward.
The area of the strap in the slot is 1*5 sq. cms., so that the resistance of 1 metre of
the conductor is 0-000115 cold. We have, therefore,

0-000115 X 1-16x7102x6 = 400 watts per metre.
The area of the insulation is about 2000 sq. cms. per metre and the thickness 015 cm.

0-0014 x<°_ 400
015 "2000'

f = 21 •5'^ C. rise of copper above iron.

It is interesting to note that with this construction we can work the copper in
the slot as high as 470 amperes per sq. cm., and yet have quite a low temperature rise.

The area of the end connectors of the rotor must be greater than the area of
the conductors on the slots, on account of the much poorer cooling conditions.
We have chosen an area of 2-25 sq. cms. The cooling takes place, partly by con-
duction of the heat through the insulation flanking the end connectors, and partly
by conduction along the connectors to the ends which are very well ventilated.
The general method of finding the temperature rise in cases of this kind is described
on page 226. This case is rather complicated by the fact that the connectors are
reduced in section at the dovetailed portion, and the flow of heat by conduction
along the copper is throttled at this point. The simplest way of getting over this
difficulty is to imagine the conductors are not reduced in section, but that they
are lengthened instead. It will be seen that both 7^, the current density, and x,
the length of the conductor, enter into the equation

Tx = T^ cos (4-71 X 10-*^ xldxx)

in such a way that to multiply Id by any constant has the same eflect as multi-
plying X by the same constant.

For instance, on the machine under consideration, in the part 2 cms. long,
where the cross-section is reduced to one-third, and the current density is increased

400

DYNAMO-ELECTRIC MACHINERY

a
s

>
cl

a
&

S3
■•»

■<

■5:

-a

o
o

►

•5

!<•
M

— i-

402 DYNAMO-ELECTRIC MACHINERY

to three times, the temperature fall will be the same as if the section had not been
reduced, but instead, the part in question were made 6 cms. long. If we take
into account the bevelled part, we find that the effect of the whole reduction in
section is the same as adding 4*5 cms. to the length. This makes the strap, which
we may judge to be the hottest, about 37 cms. from its centre point to the place
where there is presented a large cooling surface.

The watts lost in the 33 end-connectors for one pole, at a current of 710 amperes,
are 960 watts. From this we must subtract the watts dissipated by conduction
through the insulation. The cooling surface of 2400 sq. cms. at say 0-13 watt
per sq. cm. gets rid of 310 watts. 960 - 310 = 650 watts to be conducted along the

copper. 12 650. . r ^o-897,

7? = %0' •• ^^-0 82/d.

/d = 315 amperes per sq. cm. ; .*. /v = 258 amperes per sq. cm.

The law of temperature distribution is

3'x = r,„a, cos (4-71 X 10-6 X 258 X 37).

Now, the very large cooling surface exposed by the straps which pass over
from one tier to the next, and the strong blast of air blowing on this surface, wiU
keep the ends of the connectors very cool. 20° C. rise is an outside figure for the
arrangement shown. Let us say that the actual temperature of the ends is 45° C.
Add 240° (see p. 227), and we get

285 = r„^ cos 0-45,

^iii»x = 316,

316- 240=76° C. actual,

or, say, 51° C. rise in the hottest point of the connectors.

In calculating the resistance of the field winding we find that the bars have a
total length of 1320 metres, and have a resistance of 0-115 ohm per 1000 metres,
while the end connectors have a total length of 349 metres of a conductor having
a resistance of 0076 ohm per 1000 metres. The total resistance is 0-178 ohm cold,
or say, 0-21 ohm hot. To drive 710 amperes we will require 150 volts, so that the
exciter should be capable of generating about 190 volts to deal with over loads.

The working out of the efficiency will be easily followed from the calculation
sheet.

TWO-POLE TURBO-GENERATORS.

In order to get the high steam economy which is only possible at very high speeds,
the tendency is to build larger and larger units running at 3000 B.P.M. Generators
of 50 cycles, having an output as high as 5000 K.V.A., are now run at this speed.
Such a high speed does not lead to economy in the generator itself, because the
windage losses are high and the cost of construction is greater than for a four-pole
generator of half the speed. These disadvantages, however, are outweighed by the
advantages to be gained in the steam turbine. It has therefore been necessary
to overcome the inherent difficulties in building a two-pole turbo field-magnet^
and this has been satisfactorily accomplished by the constructions shown at the
beginning of this chapter.

ALTERNATING-CURRENT TURBO-GENERATORS 403

In order to get as high an output as possible from a machine of limited diameter
and length, these high-speed, two-pole machines are usually made with a rather
low ratio of field ampere-turns to armature ampere-turns, so that the regulation
is very poor. Automatic regulators are therefore commonly used in conjunction^
with them to keep the voltage constant. Very often no guarantee is given as t0|||
inherent regulation, but an automatic regulator is supplied which will hold the
voltage within 1 or 2 per cent, under normal working conditions.

For the 2500 K.V.A., 2-pole, 50-cycle generator, particulars of which are given
on the design sheet, page 406, we have chosen a rotor cut out of a solid steel forging,
because this construction enables us to make the critical speed at which the rotor
begins to whip, higher than the running speed. The performance specification might
be worded as in Specification No. 6.

404

DYNAMO-ELECTRIC MACHINERY

SPECIFICATION No. 6.

2600 K.V.A. THREE-PHASE GENERATOK TO BE DRIVEN BY A

STEAM TURBINE AT 3000 RP.M.

Clause as to General Conditions, see Clauses 1, 21, 170.

Extent of work.

80. The work includes the supply, delivery, erection and
setting to work at , of a turbo-generator

and exciter, together with automatic regulating gear. The
plant shall have the following characteristics :

Charaoteristios
of Generator.

Normal output
Power factor of load
Number of Phase
Normal voltage
Voltage variation
Amperes per phase
Speed
Frequency
Regulation

2500 K.v.A. or 2000 K.w.

0-8.

550.

520 to 570.

2620.

3000 revs, per minute.

50 cycles per second.

The generator or its exciter shall
be controlled by an auto-
matic regulator, which shall
keep the voltage constant
within 1 per cent, when a load
of 200 K.w. at 0*8 power factor
shall be thrown on or off the
generator. This regulator
shall be supplied * under the
contratjt for the supply of the
generator, and shall be in-
cluded in the price.

3300 amperes per phase at 550
volts with power factor be-
tween 0*9 and unity.
Exciting voltage 110.

Temperature rise after ] 40° C. by thermometer.
6 hours full load J55° C. by resistance.

Temperature rise after 1 55° C. by thermometer.
2 hours over load J 70° C. by resistance.

* In some oases the purchaser will akeady have a regulator installed. In these cases
particulars should be given of the type and arrangements made for including the new
exciter in the regulating scheme.

Over load

ALTERNATING-CURRENT TURBO-GENERATORS 405

81. The generator is intended to supply power to two Nature of load.
cotton factories situated at , and to

three other factories at a distance of about 1 mile. Some
1250 K.W., taken from the generator at 550 volts, will be
transformed up to 3000 volts for transmission by underground
mains to the three distant factories ; part will be consumed
without transformation, on motors varying in size from
5 H.p. to 100 H.P., and another part, about 100 k.w., will be
transformed by static balancers to 120 volts for lighting.
This Ughting load will be distributed as evenly as may be
between phases, but the phases may be sometimes sUghtly
out of balance. The generator must be suitable in every way
for this class of work.

82. The revolving parts of the generator and exciter shall cnticai speed.
be so constructed that the critical speed is not less than 3600

revs, per minute.

83. At the normal speed of 3000 revs, per minute the rotor Factor of
shall have a calculated factor of safety on every part of not ^^*^'
less than four. The revolving part shall, before leaving the
Contractor's works, be run at a speed of 3300 revs, per minute,
without showing signs of movement of the component parts
relatively to one another.

Here may follow Clauses Nos. 5, 6, 8 or its equivalent (see Clauses
55 to 59), 10, 11, 12, 13, 14, 15, 16, 17, 18, 19, 20, 23, 26, 27, 60, 61,
64, 66, 68, 69, 70, 73, 74, or such of them as are suitable for the case.

CALCULATION OF A 50-CYCLE 2500 K.V.A. TURBO-GENERATOR

RUNNING AT 3000 R.P.M.

Diameter of rotor. The considerations which determine the diameter of the
rotor are as follows. The smaller the diameter the less will be the centrifugal forces
and the less will be the windage. On the other hand, a small diameter may necessi-
tate a great axial length in order to get the required output, and a great length
makes it difficult to give to the rotor sufficient lateral stifEness. The critical speed
at which the rotor begins to whip depends upon the stiffness of the rotor regarded
as a beam supported at its two bearings. The critical speed in revolutions per
minute is equal to

X (^1^1 + ^2^2 + ^3^3 + etc.)

60 /32-2x 12:
2W (/r,y? +

where TFj, TFji e^^v *re the weights (in lbs.) of various convenient sections of the
rotor and yj, y^, etc., are the deflections (in inches) of the centres of those sections
produced by the action of gravity as the rotor is held horizontally on its bearings.

406

DYNAMO-ELECTRIC MACHINERY

ALTERNATING-CURRENT TURBOGENERATORS

407

The deflection of the rotor can be worked out by the well-known graphical
method. Where the rotor consists of steel punchings threaded on a shaft it is
found in practice that the amount of stiflness aflorded by these punchings, even
when very firmly bolted together, is generally very small, and may be neglected
in comparison with the stiflness of a strong shaft. The punchings, however, absorb
a considerable amount of the energy of the whipping action, and enable a rotor
which is not very badly out of balance to run through the critical speed without

FIQ. 877.

excessive vibration. Where the rotor is cut out of a solid steel forging the critical
speed can be pre-determined with greater accuracy than where it consists of various
parts pressed or shrunk on to the shaft.

We choose a diameter which by trial gives us the required output with an axial
length which permits of the required stiflness of shaft. If the diameter is about one-
half the axial length of the iron, the proportions are generally good mechanically,
and economical electrically, for two-pole turbo-alternatives. The general proportions
of the machine are the same as those shown in Figs. 376 and 377, which show a
50-cycle 1500 K.W, generator built by the A.E.6. for a speed of 3000 r.p.m. Figs.
354 to 357 (page 369) show the method of constructing the rotor of this machine.

It will be found that for these high-speed two-pole turbo-generators with poor
regulating qualities we can take an output coefficient between 500,000 and 600,000,

408

DYNAMO-ELECTRIC MACHINERY

Date^^/eAipA? TjptTur6o.AC.CMn..

KVaJ^SOO: PF.*.<8: Phue3 : Volts S50
H.P. Amps p. cond. /3/0 Amps p. br.

arm.

. J?.. J\>les . -^ . . - Elec Spec 6f ,
Amps prr ter 2620.. Cycles s5(? ; RPU3000. Rotor Amps
Temp rue 40**.C Re^latwn Overload «?5'% ^/?A-5

Customer

Orde' No.

Quot No , Perf Spec

Fly-wheel effect

poss A( B.

K. -3^2 . 550 -vous '392 50 ^ 16 - I 55

; poss laZa

i.z. ^7200

Arm. A.T p. pole.

'illy

. . Ctrcum.

/a 300

197

K.V. A.

:z 5-6x10^

Max. Fid AT 26,000

Armature.

Dia Outs.--.

Dia Ins

Gross Length
Air Vents

Stat.

JZZ

o
o

e$j^'

2L . -■_] -^5''

Opening Min23^^ Mean !4^^^ IS^.OZL

Air Velocity ^2 m^fi^SeC

Net Length j&S x 89 5^3
Depth b. Slots - . J^ - , ^

sertion 2LZ^^\'o\ 6:^..\ia^

Flux Density /^^ 7Q^

Los&:fl!6.p.cu.i:/a.TotaI 3ft.fi^-<

Buried Cu ^23(2^Total 4i^QQ^X0Q

GapArea7>«0/?P.VVU ZZ3.flO.

VentArea45a<^?a^Wts 32J0QO.

Outs. Area4MQ^. Wts

»-

No of Segs I 6_lMn.Circ.
No of Slots 3fiix2^ =
K.

Section Teeth .
Volume Teeth.
Flux Density,

.6500.

36-5

J2SiS.

— ^/0L600

\ Loss '/ - p. cu Cifla-Total

m^ooo

l^^500_

c
o
o

Weight of Iron.

-Throw

Star

Cond. p Slot

Total Conds ^^^

Size of Cond IlS-xJlS. 3l5SSfSim^

Amp. p. sq.-CJTl '^

^0300

5500 \kUQ^fl

IbLl^

talSJ^

Length in Slots -JI3 .
Length outside i^ZSum
Total Length

J3.
3i
J313

262
SSm

wt. of i.oooaS/^^Totai ' J3QOj!ci/o4r^

Res. p 1. 000 '^A^Total

Watts p

Surface p.

Watts p. Sq

56 Slofs

I ^ f'9^

32a^ if 42 Slots

Field

Rotor.

Dia Bore

i Total Air Gap

Gap Co-elf K,

Pole Pitch Pole

Kr -

Arc

J3.

2-25

I OS

Flux per Polc<^?^l5 ^^'.

Leakage n i^lO\\6j'lQ^ 53'S^O^

62^ :

Area SSM Flux density _ fS.^OO-,
TJnbalanced Pull L

No ofScg.

_7. mnCiic

No of Slots

« .. x/-p=

Vents 2/ 1 JW?

K,3 *B

-Section

/28'^

iS7

m 22.

4^ 4. . 77

Weight of Iron

A.T. p Pole n. Load
AT.p Polef Load
Surface .

Surface p Watt

V R

I R.

Amps.

No. of Turns

Mean I Turn

Total Length

Resistance

Shunt, t

itQQo:.

F.L

25%

2&.SOO 29.000

3JL

4^00_
4-0-5

I5JOO

97 117

M>J>.

/44

l^ao_

250cm ISO cm^ 380

7_20jtt±322jrL

- '395co/d'fr7Shot

Res per I 000 |" 3&^nd V^

Size of Cond , ' 305xf53nc/y3O5XI 9

JA 1.

per Slot

Conds

Total

Length .

Wt per i.ooo-
ToUl Wt _

Watts per Sq

Star or Mesh \

Paths in parallell

I092
4^5

>-£

Magnetization Curve.

Core

Stator Teeth
Rotor Teeth

Gap

Pole Body -
Yoke

Section. Lsngth

2120 €3 6S2O0 3 5 _2ZQ UAQO . 5_ . 320 _ fUQO £. 380.

/aeOO ^ i^Oa.^^:^ 36 ia.,6pO_J§ _ e^. 15200 -2J .. -^4L Volts p. Bar

lazo^ 9

^/OC 2-25

600^\/ojis.

. iA.T.PC/»^A.T.

4(7

/2300\
"ZOO

ISese

j5l5jOvj)its.

B. l/LT.pc/flAT

/3700

yoQc

-20a-

f??Ka

CFriCIENCY

ili load., Full. I f

Friction and W.

Iron Loss

Field Loss

Arm. &c. rR_
Brush Loss _

■1 4^.!-^6_.4g

' 4.7 I 4-7 I 47

21^^ fS

13 fi.

/26\Tfe

Output -
Input —
Efficiency

^'5_
iSJL

/09\ tOS^ 99

2506 200d{ 1 500^000 '500
262G 2tf6
952 9^5

1609 1/03 ■ $99
9M 90 5^ 35 \ S,;

SCO

aitxrz^

S7S Vo\is

B. Ll.i-cmLl

.^. foo . »-^o^_i6 M^
fssoo 7,350 .. . .i^JZQi

1?096\ r^.ooo

fSOO

r-f-f'T'i

Connmutator.

Dia_
Bars

3peed _

Brs. p. Arm
Size of Brs
Amps p sq
Brush Loss
Watts p Sq

Mag Cur

Perm Stat. Slot
,. Rot Slot X
.. Zig-zag
X

Loss Cur

2 X

177
End

X
X X

Ampb . Tot
: X. =

= +

Imp v" +

Sh cir Cur

Starting Torque
Max Torque _

Max. H.P

SUp

Power Factor

ALTERNATING-CURRENT TURBO-GENERATORS 409

the dimensions being in centimetres. An internal diameter of stator of 63-5 cms.
gives a rotor diameter of 59 cms., and an axial length of 115 cms. These proportions
are suitable, the output coefficient being 560,000.

A section of the iron of the stator is given in Fig. 240. Particulars of the iron
loss and windage loss, as determined by experiment, are given on page 244, and the
distribution of temperature with various amounts of cooling air are given in Figs.
241 to 247. The rating of the machine upon which these tests were carried out
was 1870 K.W., but that rating was fixed in order to meet a certain regulation
guarantee. The frame can be rated at 2000 K.w. for a poorer regulation with the
iron worked at exactly the same state of saturation.

Particulars of the windings on stator and rotor are given in the calculation
sheet on page 408, and method of calculating the various quantities will be easily
understood from the description of the method given on pages 316 and 332. It is
therefore unnecessary here to go through the sheet in detail, but the reader will
be interested in comparing the results arrived at by this method of calculation
with the actual results experimentally obtained. The rotor winding consists of
concentric coils of the type shown in Fig. 359, but the parts of the coils lying outside
the slots have a section of copper 0305 x 1-9 cms., while inside the slots the section
is 0 305 X 15 cms. This helps doubly in keeping down the temperature of the end
connections. It gives a lower current density, and it gives a great section of copper
for the conduction of the heat to the straight parts of the coils, where most of the
cooling surface is (see page 225).

25-CYCLE TURBO-GENERATORS.

A two-pole machine to generate at 25 cycles cannot have a speed * higher than
1500 R.P.M. This is a drawback from the turbine builder's point of view, and is one
of the reasons why 25-cycle turbo-generators are not often btdlt in small sizes.
Even for large sizes, where 1500 r.p.m. is quite economical for the steam turbine, the
two-pole generator is much more costly than a four-pole generator of the same
output. On account of the lower speed the diameter can be increased, so that
outputs up to 25,000 K.W. or higher become possible. The very bulky end connec-
tions on these large two-pole machines make a very undesirable feature. The
general proportions of a 25-cycle turbo-generator for 1500 R.P.M. will be seen from
Fig. 378, which shows a 2500 k.v.a. 25-cycle turbo-generator built by the Oerlikon
Company.

* Certain methodfl of construction have been suggested for enabling 25-cycle generators to
be run at speeds higher than 1500 b.p.m. In some of these the field of the rotor is a polyphase
field, which rotates backwards relatively to the rotor iron. In another ingenious suggestion the
poles on the rotor are distributed like the thread of a screw around the rotor surface, and
the speed of moyement of the pole relatively to the stator conductors can be made as slow as
desired by making the pitch of the screw very small. These methods have not come into
general use.

I* X-

^^^H _ J WW ^ — ■ T'-f

i *- I — ^ V ! ■ ^ t

Ifff

-wr

TET

I _^

r.^

- rJ

I- ^ — -r^ — f -"^ 4

r-izj

rr-zn

, -J

>»

A
I

•

ALTERNATING-CURRENT TURBO-GENERATORS 411

SINGLE-PHASE GENERATORS.

Single-phase turbo-generators are sometimes now required for central stations
which adopted a pingle-phase system in the early days of electric supply, and
which have not yet changed over to polyphase. A certain number of low-frequency,
single-phase generators are manufactured for single-phase traction, though where
possible it is more economical to build a polyphase generator, and arrange for the
different phases to be supplied to different parts of the system. Single-phase
windings have been considered in Chapter YI. The most common practice is to
take an ordinary three-phase armature and put the winding in two-thirds of the slots.
Sometimes, for convenience in getting the right number of conductors, one will
use rather more or rather less than two-thirds of the slots. It is not well to use
much more than the two-thirds, or we shall get some coils which enclose only a
small fraction of the total flux, and thus employ a large weight of copper for the
output. The mechanical arrangement of the end connectors of the armature is,
of course, much simpler than on three-phase machines, and the overhang of the
coils is reduced.

The aimatnre Teaction of these machines is pulsating in character, and gives
rise to pulsations in the field-flux, which may cause serious heating of the field-
magnet, if proper precautions are not taken to prevent it. The most common plan
is to provide the field-magnet with a damper or amortisseiir, the eddy currents
in which oppose any change in the value of the flux. This damper can be con-
veniently made on cylindrical rouors, by using copper wedges in the tops of the
slots, and connecting electrically and mechanically with conducting end rings, so
as to form a squirrel-cage winding. In calculating the cross-section of copper
to be used in this squirrel-cage, we must remember that a single-phase armature
reaction may be regarded as due to the sum of two vectors rotating in opposite
directions. Each of these vectors represents a number of ampere-wires equal to
one-half of the total ampere-wires on the armature. The vector, which rotates
in the same direction and at the same speed as the field-magnet, does not produce
any pulsation, but only a steady distortion, just as a polyphase reaction (see page
278). The vector which revolves in the opposite direction produces an eddy current
in the squirrel-cage winding of double frequency. The number of ampere-wires
in the phase-band of eddy current is equal to one-half the phase band of current
on the armature. We must therefore provide a cross-section of copper in the
damper sufficient to carry one-half of the ampere-wires on the armature. At the
same time we must remember that the high frequency of the damper current will
produce eddy current losses in any solid metal parts enclosed in the magnetic
circuit of the damper winding, and accordingly arrange all surrounding parts so
that they are either of such good conductivity that the eddy current can flow
without causing excessive loss or of such high resistance that no appreciable loss
can occur in them.

If we take an ordinary star-connected, three-phase generator, having terminals
Ay B and C, and load it as a single-phase machine by putting a load across the
terminals A and B, we will find that one kilowatt of single-phase load will produce

412 DYNAMO-ELECTRIC MACHINERY

about 1 -35 times as much reaction on the field-magnet as one kilowatt of three-
phase load. As the output of the generator is usually limited by the output of the
field-magnet, we may say that as a single-phase generator a frame will only carry
about 0-74 of the load it would carry as a polyphase generator.

The reader will find useful data relating to single-phase generators in the articles *
quoted below.

* "Modem Development in Single-Pha«e Generators,** W. L. Waters, Amer, I.E.E.f Proc.
27, p. 679, 1908; "Comparative Capacities of Alternators for Polyphase and Single-Phase
Currents,** Elec. Joum., 8, p. 672, 1911.

CHAPTER XVI.

INDUCTION MOTORS.

We shall assume that the reader is familiar with the general theory of the induction
motor and the use of the Heyland circle diagram. Different writers give the circle
diagram of the induction motor in different forms. It is therefore convenient tx)

Fio. 400. — Circle diagram of induction motor.

reproduce here a form'^ which is found to be very convenient in workshop use, and
to give results which check sufficiently well with those obtained in practice.

*See Karapetoff's Experimented Electrical Engineering^ vol. 2, page 166 (Chapman &
Hall, LondoD ; John Wiley & Sons, New York) ; Cramp and Smith, Vector Diagrams (Longmans) ;
** Graphical Treatment of the Rotating f^eld," B. E. Hellmund, Amer. I.E,E., Proc. 27, p. 927,
1908 ; " Circle Diagram for the Induction Motor," J. Yemaux, 8oc. Beige. Med., Boll. 27,
p. 246, 1910 ; " Circle Diagram of the Induction Motor," W. Petersen, EWdrotech. Zeitschr.,
31, p. 328, 1910 ; " Polypha^ Induction Motor, Circle Diagram of,*' K. Krug, Elek. u. Maachinen-
ban, 28, p. 1047, 1910 ; ** Circle Dia^m for S-phase Induction Machines,^' T. F. Wall, LE.E.
Joum., 48, p. 499, 1912 ; ** Vector Diagram of the Induction Motor on an Experimental Basis,*'
L. Dreyfiis, Archiv.f. EleHrot., 1, p. 124, 1912 ; " Simple Graphical Construction for Determining
the Efficiency of a Polyphase Asynchronous Motor from the Current (Circle) Diagram,** J.
Nioolson, Joum. LE.E., 49, p. 297, 1912.

414 DYNAMO-ELECTRIC MACHINERY

When the no-load data of the motor are available we can construct the Heyland
diagram shown in Fig. 400 as follows :

Choose a current scale (say, 1 mm. = 10 amperes). Let the lines on the diagram
represent amperes per phase to this scale.

ON represents the no-load current per phase.

N^N „ wattful current per phase supplying the no-load losses.

ON' „ true magnetizing current per phase.

O'/S ,, current per phase on short circuit.

SF „ wattful current per phase supplying the losses on short

circuit.
TF „ wattful current per phase supplying the losses in stator

on short circuit.

ST „ wattful current per phase supplying the losses in rotor

on short circuit.
OP „ stator current on normal load.

NP „ rotor current on normal load x —2.

PX ,, wattful current supplying useful work.

PY „ wattful current supplying power to rotor.

PZ „ wattful current supplying power to stator and rotor.

XY „ wattful current supplying rotor losses.

YZ „ wattful current supplying stator losses.

<t> f, angle of lag neglecting stator losses.

<l>' „ angle of lag after approximate correction for stator losses.

UW „ wattful current per phase supplying the maximum power

of the motor.
QR f, wattful current per phase supplying the maximum torque

to the motor.
The slip is given by the ratio XY : PY.

To convert any vertical line into watts, multiply the number of amperes per
phase by the line voltage and by 1 -73.

To convert a vertical line into torque in lbs. at a foot radius. Multiply the number
of amperes (obtained by scaling it off) by

Line volts x 1-73 x 33,000
746 X syn. Rpm x 27r

The usual method of procedure is to set off the no-load current ON to scale.
To find its position with respect to OE we may either set off EONy the angle of lag
at no load, or we may set off N'N, the wattful current at no load (see page 420).
Then set off O'/S, the short-circuit current. The point 0' is chosen, so that

7vifT= — 7 — ' 3 . Usually it is sufficient to put 0' half-way between 0 and

O N rotor impedance "^ r ^

N. SF can be calculated by dividing the watts lost on short circuit by the line

volts and by 1-73. Thus we obtain the load line NS, To get the centre of the

semi-circle, we bisect NS in V and draw VC at right angles. Where this cuts the

horizontal line NF is the centre C. We can then draw the semi-circle through

INDUCTION MOTORS 416

N and S. This gives us the locus of the point P for normal loads. For loads heavy
enough to cause saturation of iron along the leakage paths, the point P moves on
a rather wider curve, as shown by the dotted curve in Fig. 415. For normal loads,
and even up to the maximum output, the semi-circle NPU gives the locus of P
with sufficient accuracy for practical purposes. It is usual to take the angle <f>
as the angle of lag of the stator current behind the voltage, but this is only right
when the stator-resistance drop and the change in the value of ON can be neglected.
An approximate method of allowing for the effect on the power factor of the stator
resistance and the change in the value of ON is to shift the position of the origin
0 and make it travel around the little semi-circle 00p0\ as P travels around its
semi-circle. Thus, if we take the origin at Op instead of at 0, we get the angle
4>' as the angle of lag, and this is rather smaller than 0. In actual practice, however,
the power factor depends somewhat on the wave-form, and as this is generally
unknown when a motor is being sold, it is safer to base guarantees of power factor
on the angle <^, and keep in hand any advantage that may afterwards be derived
from the fact that </>' is smaller.

Any vertical line drawn from P and cutting NS, the load line, gives by its
intercept (such as PX) the wattful current per phase, supplying the output of the
rotor. It is thus proportional to the output of the motor, and to get the output
in watts it is only necessary to multiply the number of amperes represented by the
vertical intercept by the voltage and by 1-73.

The maximum output is obtained from the vertical line VW, drawn from V,
where the tangent parallel to NS touches the semi-circle.

To get the torque line we must divide FS into two parts, such that

STr^^^xNS^
TF^ r, X O'S^ '

when r2.i and r^ are obtained as shown on page 428. Any vertical line drawn from
P cutting the torque line NT gives, by its intercept (such as PY), the wattful
current per phase supplying the input to the rotor. As the torque multiplied by the
synchronous speed gives us the input into the rotor, we can obtain the torque in
lbs. at a foot radius by multiplying by the constant given on page 414. The
maximum torque is obtained by drawing a vertical QR from the point Q, where a
tangent parallel to NT touches the semi-circle.

Circle diagrams drawn to scale are worked out on pages 457 and 474.

In working out a circle diagram we may be able to start with data given by
experiments on the motor in question, or we may have to start with the particulars
of the design, and deduce the leakage flux and magnetizing current, etc., from
the dimensions. When the no-load and short-circuit data are given, the working
out of the power factor and other particulars of performance from the circle diagram
is a comparatively simple matter, and is indicated in the method followed in the
examples given below.

When the no-load data are not given and have to be deduced from the dimensions,
the calculations are somewhat more lengthy, but they can be shortened by the
judicious use of formulae founded on practical results. The principal quantities
to be determined are the magnetizing current and the leakage flux.

416 DYNAMO-ELECTRIC MACHINERY

DETERMINATION OF THE MAGNETIZING CURRENT OF AN

INDUCTION MOTOR

Following the general method adopted throughout this book, we concern
ourselves first with the maximum flux-density in the gap. This will usually be
found to be much lower in induction motors than in a.c. generators, because of
the importance of keeping down the magnetizing current. If too high a flux-density
in the gap were chosen, not only would the magnetomotive force on the gap be
great, but the teeth, being necessarily of wide section to carry the heavy flux,
would leave little room for copper, and thus the number of turns per pole would
be few and the exciting amperes in consequence great. In an A.c. generator we
choose a low number of turns per pole on the armature to improve the regulation.
In an induction motor we choose a high number of turns per pole, in order to keep
down the magnetizing current. The increasing of the ampere-turns on the armature
of an induction motor must not, however, be carried too far, because the more we
reduce the flux per pole the greater we make the ratio between the leakage flux
and the working flux, and the smaller we make the diameter of the main circle of
our diagram.

The flux-density in the gap. The maximum flux-density found in induction
motors usually lies between 5000 and 7000 lines per sq. cm. B = 6000 is a common
figure. In low-voltage motors of large size it will be more, and in high-voltage
motors it will be less. In motors designed to be used in conjunction with a phase
advancer, the flux in the air-gap may be carried to as high a figure as 9500. The
reader will see from design sheets on pages 448 and 471 the general considerations
which fix the density in the gap. Often in high- voltage motors the amount of room
taken up by the stator coils and insulation so reduces the section of the teeth as
to necessitate a rather low value for the flux-density in the gap.

Length of air-gap. The length of air-gap in induction motors is made as short
as is compatible with securing a good mechanical clearance. In very small motors,
particularly if the surfaces of the rotor and stator are ground perfectly true,
exceedingly small clearances, even down to 0*04 cm., may be employed. On large
machines the length of air-gap is generally increased. The curve in Fig. 401 shows
the relation between the length of the air-gap and the diameter of rotor according
to good practice, where precautions are taken to avoid distortion of the frame.
If the air-gap is made too short, the unbalanced magnetic pull due to extremely
small accidental displacements may be excessive, and by causing a further dis-
placement may bring the rotor in contact with the stator. Where it is desired
to keep the power factor of a large induction motor with numerous poles as high as
possible, and therefore the magnetizing current as low as possible, the designer is
tempted to reduce the air-gap to the smallest permissible figure. For this purpose
he arranges the stiffness of the frame and shaft so that they will withstand .a
heavy unbalanced magnetic pull without undue distortion. One method of greatly
reducing the unbalanced magnetic pull is to connect the two halves of the stator
winding in parallel, each half of the winding occupying coils on opposite sides of
the diameter, in the manner shown in Fig. 409. In this case the division of the
two halves of each phase should take place about diameters placed at angles of

INDUCTION MOTOBS

417

60° to one another. It is not then possible for the magnetic flux on one side of
one of these diameters to be much greater than on the other, because the greater
flux would produce a greater back electromotive force and keep down the
magnetizing current on the side which, by reason of its short air-gap, might
otherwise tend to have an excessive magnetic flux.

•35

9»

1-30

fe-25

1.
O

o
■

2-05

50 100 150 200 250 300 350

Diameter. of Rotor.in Centimetres.

Fio. 401. — ^The length of air-gap in InductioD motors.

400 450

The calculation of ampere-turns on the gap. The general method of calculating
the ampere-turns on the air-gap, given on page 66, is applicable to induction
motors. In general, the air-gap coefficient Kg of an induction motor is much
higher than for an A.c. generator, on account of the greater ratio of width of slot
to length of gap.

£x AMPLE 48. Take the dimensions of slot, gap and ventilating duct from the calculation
form on page 448. We have

-=p-—=l'o and —=s-5 =0*115.
y 0-2 p 2-6

From Fig. 37, page 67, we have the contraction ratio for stator slot 1*03. Similarly for
rotor slots it is 1 *04. Next, for the ventilating ducts

« 0-8 .^ , « 7x0-8 ^ ,„
- = irs=4-0 and — = — 7^^ = 0*12.
g 0-2 p, 47

From Fig. 36 the contraction ratio for the ventilating ducts on the stator is 1*05. It is the
same for the rotor ducts. Taking the product of all these, we have

103 X 1-04 X 105 X 105= 118 = iry

for the gap of the induction motor.

The ampere-turns on the gap are equal to

6700 X 0-2 X 1-18 X 0-796= 1070.
w.M. 2 D

418

DYNAMO-ELECTRIC MACHINERY

The calculation of ampere-tums on the teeth. This is carried out in the same
way as described on page 73. In many cases the ratio Kg (see page 71) in induction
motors is fairly high on account of the small section of the teeth ; and where high
saturations are used it is desirable to have recourse to Fig. 47 in order to find the
actual ampere-turns on the teeth, because the apparent flux-density difEers appreci-
ably from the actual flux-density.

The permissible flux-density in the teeth is limited in 50-cycle motors by the
permissible iron loss per cu. cm. of tooth. It may be between 16J500 and 17,500,
depending upon the cooling conditions. At low frequencies the density is limited
by the number of ampere-turns that may be applied to the teeth. Too high a
density will require too great a magnetizing current and spoil the power factor
of the motor. Densities of 18,500 to 20,000 lines per sq. cm. are not imconimon

5000

iooo

3000

2000

1000

•^

Bi/

.A A ^

3800

'3/50

////

soot

//a

"990

1//

.. .

1 t

;

200 900 600 800 WOO AW 0' 30' 00*

Fio. 402. — Method of finding the average field-form of an induction motor.

OCT

in 25-cycle motors. Where a phase advancer is used in conjunction with the motor,
and the frequency is not so high as to make the iron loss excessive, densities as high
as 21,000 lines per sq. cm. are permissible.

Strictly speaking, the maximum density in the teeth can only be ascertained
at any particular voltage on the stator winding by plotting the field-form of the
motor, and from it the values of the e.m.f., as we have shown on page 32. In
actual practice, the ampere-turns required for the teeth are known from experi-
ments upon the particular stampings in question, for every state of saturation of
the frame, so that no elaborate calculation is necessary. In cases where the ampere-
turns are not already known from trial, and where the saturation is not very exces-
sive, say, not higher than 17,000 lines per sq. cm., we may assume that the maximum
density in the teeth is what it would be with a sine distribution of flux, and make
our calculations accordingly without introducing very great error.

Where the saturation is very high it is possible to ascertain the approximate
shape of the field-form by the method given on page 21, aided by a curve giving

INDUCrriON MOTORS 419

the ampere-turns on gap and teeth for each value of the flux-density on the gap.
The method of plotting such a curve is given on page 78.

In Fig. 402 the curve Bz on the left gives the ampere-tums (a.w.) on the teeth
of a 3000- volt induction motor for different values of B in the gap. The straight
line Bi gives the ampere-tums in the gap. The combination of these gives the
ampere-tums on teeth and gap. Taking now a certain magnetizing current
passing through a distributed winding, giving, say, a total of 700 ampere-turns per
pole, we can, in the manner shown in connection with Fig. 15, aided by our combined
tooth and gap saturation curve, plot and approximate fleld-form shown by the
lower full line curve on the right. From this field-form we can calculate the generated
E.M.F. (see page 32). Say that this is 2940 volts. Now plot another field-form
for a higher total number of ampere-turns, say 800, and calculate the e.m.f.
generated by that field-form. Say that this is 3150. Then the ampere-turns
required for the normal voltage of the machine will, for small variations, be almost
in proportion, so that for 3000 volts they will be about 730

The amount of space taken up by the rotor copper and insulation is usually
much smaUer than that required for the stator copper and insulation. It thus
comes about that there is usually a more liberal allowance of iron in the rotor
teeth, so that the number of ampere-turns on these teeth is often small.

Ampere-tums on the core. In low-frequency machines which have a fairly great
pole pitch, and in which the flux-density in the iron is carried up to a fairly high
point, some allowance must be made for the ampere-tums required on the cores
behind the slots. The amount is small in comparison with the other ampere-tums,
and therefore no time need be wasted in making an accurate calculation. It is
sufficient to assume that the ampere-tums on the core are equal to the ampere-
tums that would be required on a core length of one-third of the pole pitch, in which
the flux-density is equal to the maximum density found in the core of the machine
in question.

Example 49. Id the motor illustrated in Fig. 407, the pole pitch is 31*2 cms. The maxi-
mum flux -density in the core is 8450 lines per sq. cm. Find approximately the ampere-tums
per pole required for the core. Take the eflfective length one-third of the pole pitch, say
10*5 cms. For a flux-density of 84o0, we require about 3 ampere-tums per cm. 10*5 x 3 = 32
ampere-turns per pole on the core. These are so low that in practice they could be neglected,
because they are smaller than the errors coming into other parts of the calculation.

Suppose now that the motor was working at 25 cycles with a flux -density in the core of
16,000. The ampere-tums on the core would then be 8x30=240.

In the calculation sheet given on page 448 will be found a table of the ampere-
turns required on various parts of the magnetic circuit when the motor is operating
at 3000 volts.

Magnetizixig current. After we have found the total number of ampere-tums
required to produce the flux-density in the centre of the pole, it remains to cal-
culate the number of virtual amperes of magnetizing current to be supplied to each
terminal of the motor. The commonest case with which we shall have to deal
will be the case of a three-phase star-connected stator having a full pitch winding
arranged as in Fig. 110, with two, three, four or more slots per phase per pole. For
this type of winding it has been shown on page 280 that the average value of the

420 DYNAMO-ELECTRIC MACHINERY

ampere-tuniB per pole exerted by the armature is approximately equal to 0'437

A T

ImZa^ 80 that Jm = A A*>trfT » where Im stands for the virtual value of the wattless

magnetizing current. To arrive at the core loss current le, we divide the calculated
(or measured) iron loss by the voltage and by 1*73. The total magnetizing current
Inu will then be ^/^^ + i^. It can be conveniently obtained by a graphic con-
struction. Where the winding is not of the common kind presupposed here, the
safest plan is to lay out a diagram of the slots with the windings belonging to the
various phases indicated in their respective positions. Then, assuming that one
phase is at its maximum and the other two phases at one-half their maximum,
plot the magnetomotive force wave produced thereby, as shown in Fig. 15. From
this the virtual amperes per phase required to produce a certain maximum number
of ampere-turns on the centre of the pole is at once apparent. Then take two
phases at 0'866 of their maximum value, while the other phase is at zero, and
make a similar plot from which the virtual amperes per phase for the same
ampere-turns on the pole can be ascertained. A mean of the values obtained
in the two cases will give the magnetizing current with sufficient nearness for
practical purposes.

The no-load current. If the no-load losses (iron loss and friction losses) and
the magnetizing current are known, the no-load current is obtained as follows.
Let Wn be the no-load losses in watts and Et the volts at the terminals ; then

W
pr^-^-=rs= current per phase suppl3ring the no-load losses = 7n2.

Set off Im horizontally, 0N\ and Iru vertically, N'N ; then the hypotenuse
ON, is the no-load current per phase (see Fig. 400).

DETERMINATION OF THE SHORT-CIRCUIT CURRENT BY CALCULATION

FROM THE DESIGN.

If the rotor winding of an induction motor be short circuited and voltage applied
to the stator, the windings of the stator and rotor form a compound impedance
the value of which depends upon (1) the amount of magnetic flux leaking between
the primary and secondary members ; (2) the ohmic resistance of the two
windings.

The most accurate method of predetermining the short-circuit current of an
induction motor is from tests on motors built on the same or similar frames. This
is the method generally adopted in practice. The full calculation of the short-
circuit current from all the factors which influence it would be a very lengthy
matter, and at best would not be very accurate, because there are always some
factors (such, for instance, as the amount of saturation of the iron) which depend
upon accidents in the construction of individual motors. Any method of calcula-
tion from the dimensions of the motor, if it is to be of practical service, must be
fairly short. In a short method we must be content to take into account only
the most important factors, and aim not so much at an accurate determination of
the short-circuit current in any particular motor, as at an appreciation of the way

INDUCTION MOTORS 421

in which difierent factois afiect the result. The method given here enables the
designer to judge between alternative designs for the same motor, and to tell
roughly which will give the larger short-circuit current. At the same time, the
method is probably as accurate as any other method, when we take into account
the way in which indeterminate factors always influence the result.

On this subject the reader is referred to the articles *** mentioned in the
footnote.

The value of the short-circuit current depends mainly upon the ratio between
the total working flux </>/> and the leakage flux <l>i for a stator current of 1 ampere.
If the resistances of the windings could be neglected (and in practice they aflect
the result to only a very small extent), we could say that the leakage flux set up
on short circuit is great enough to generate in the stator winding as much back
E.M.F. as the total flux does at no load. If we neglect the difference in the breadth
coefficients which aflect the E.M.F. generated by the fluxes, we can say that the
leakage flux on short circuit is equal to the working flux at no load. If we further
assume that the leakage flux is proportional to the stator current, we may write
<t>i for the leakage flux per pole for one ampere in the stator, and Ia<l>i for the leakage
flux per pole for any stator current /«-

Then, if In4>i is the leakage flux at no load, and <t>p normal flux per pole,

This ratio between the no-load current and the short-circuit current, or between
the leakage flux at no load and the total flux here denoted by t is a very important
ratio, and forms the basis of the construction of the Heyland diagram. It depends
upon the ratio of the magnetic reluctance of the main magnetic circuit to the
magnetic reluctance of the leakage paths. Any change in the design of the motor
which increases the reluctance of the leakage paths or decreases the reluctance
of the main magnetic path, will decrease the value of r and increase the ratio of
the short-circuit current to the magnetizing current.

The behaviour of an induction motor when short circuited, with the rotor
locked so that it cannot revolve, is similar to the behaviour of a short-circuited
transformer having considerable magnetic leakage between the primary and
secondary coils. The main difficulty in calculating the impedance from particulars

* " Leakage Problems of Induotion Motors," R. Goldschmidt, Eketrician, 69, pp. 236, 352,
430, 507, 624, 1907-8 ; " Leakage Factor of Induction Motors," R. E. Hellmund, Elec, World,
V. 60, p. 1004, 1907 ; Elec. World, 51, p. 179, 1908 ; EUct. Rev,, N.Y., 52, p. 172, 1908 ; Ekktrot.
ZeiUchr., 30, p. 25, 1909; EUktrotech, Zeitachr., 31, ^jp, 1111 and 1140, 1910; LE.E.Joum.,45,
p. 239, 1910 ; ** Predetermination of Short-circuit Current of S-phaae Liduction Motors,** W.
Oelschl&ger, Elektrotech, ZeUachr., 28, p. 1230, 1908 ; " Determination of the Circle Coefficient
of the Induction Motor," H. M. Hobart, Elec, Rev, and West. Eleetn,, 55, p. 1073, 1909 ; " Calcula-
tion of Overhang Stray Flux in Induction Motors,** U. Kloss, Elek. u. Maachinenbau, 28, p. 53,
1910; " Leakage of Induction Motors,** W. Rogowski, Elektroi. Zeiiechr., 31, pp. 1292 and 1316,
1910 ; " Induction Motor Design Constants,** A. M. Gray, Elec. World, 58, p. 1699, 1911 ;
"Induction Motors, Reactance of,** J. Rezelman, Electrician, 66, p. 857, 1911; "Doubly-
linked Dispersion of Asynchronouslyiotors,** F. Niethammer & E. Siegel, EleHroi. u. Maschinen-
ban, 29, p. 635, 1911 ; "Experimental Determination of Leakage Factor of Transformers and
Induction Motors,** Beniachke, EleJOr. Kraflbetr. u. Baknen, 10, p. 83, 1912 ; " Air-gap Leakage
Fluxes in 2-phase Motors and in 3-pha8e Motors with 2-pha8e Rotors,** Meyer- Wulfing, Archiv.
f. EWdrct., 1 jp. 363, 1912 ; " Teste on Induction Motors designed with Deep Rotor Slots,*' L. D.
Jones, Oen. Elect. Rev., 16, p. 229, 1913.

422 DYNAMO-ELECTRIC MACHINERY

of the design lies in the estimating of the amount of magnetic leakage. Most
writers divide the leakage flux into four parts :

(1) The leakage across the stator slots.

(2) The leakage across the rotor slots.

(3) The zig-zag leakage.

(4) The leakage around the ends of the coils both on rotor and stator where

they project from the iron.

In addition to these there is a certain amount of leakage which interlinks with
both stator and rotor windings where the m.m.f. of one does not balance the
M.M.F. of the other.*

Slot leakage. The calculation of the amount of effective leakage across the
slots is most easily carried out by means of the formula

Xrf-=

1 hr
3 V

where he is the depth of the slot after a deduction has been made for the thickness
of the insulation between the copper and the bottom of the slot, and h is the breadth
of the slot. By k^ we denote the lines across the slot per cm. of axial length of slot
for unit magnetomotive force. To this must be added the leakage across the
mouth of the slot. Whether the slot is open or semi-closed the permeance across
the mouth of the slot can be found from Fig. 54 (p. 81). This figure is constructed
so that a designer can tell at once from inspection the effect of changes in the shape
of the lips upon the permeance. The shape of the lip is indicated by shading,
as shown in the figure, and the shading may extend either to the line OA, as shown,
or to the line DO, or to the line 0-25. The position of the small face P may be

varied, so that the fraction °^?? , ^^ ? ^ has any value between zero and 1. At

width of slot

whatever point we choose to draw P, it is only necessary to continue up the vertical
line from P shown in the figure until it cuts one of the curves C, <4' or B^, corre-
sponding to the depth of the lip, and we can at once read off the permeance A^ per
cm. of axial length of slot. For example, in Fig. 54, the lip is supposed to be of

the shape indicated by the shading, the value of ^^^}^ ^\ f ^^ being 0-375. If
^ ^ ^' width of slot *

we carry up the perpendicular from P to the curve A', we find that the permeance

in c.G.s. lines per cm. length of iron is 0-98. Had the lip been of a deeper design,

so as to extend up to the dotted line DC, we should have carried our perpendicular

up to the dotted curve C, and the permeance would then be found to be 1-2.

If the lip is of a special shape, or has the angle of one of its faces different from

that shown in the figure, it is easy to sketch on our figure a lip having the same

permeance and having face angles enabling Fig. 54 to be instantly applied.

Example 50. Take the stator slot belonging to the 1500 h.p. motor shown in Fig. 408a.
Here the value of he is 3 "7 and 6 = 1 '5.

* '' Leakage in Induction Motors," W. Rogowski & K. Simons, EUktrot. ZeiUchr., 30, pp. 219
and 254, 1909.

INDUCTION MOTORS 423

Now the ratio -r^ = p^=0*2, and the shape of the lip is such as to be bounded by the line

OA in Fig. 54. Therefore

X« = M3, Xd + X«=l'93.

When calculating the leakage due to the rotor slot, it is convenient to multiply
the sum of A^ + X^ obtained in the way shown in the last example by the ratio

No. of stator slots r^^^ enables the result to be added directly to the stator
No. of rotor slots

permeance, and the total leakage can be calculated from ampere wires in the

stator slot.

Example 51. Take the rotor slot belonging to the 1500 h.p. motor (Fig. 408). Here the
value of Ae = 3-6 cms. and 6=0'96.

X 1 3-6 , „-
^'' = 3 0^=^'2^-

The ratio ^=^=0*31. Therefore Xm=l-2 and \,-f-\«=2-4o.

Now there are 288 slots in the stator and 360 in the rotor, so that the total permeance
of stator and rotor slots is

l*»3 + ^x2-45 = 3-89 (seep. 448).

Zig-zag leakage. There has been a great deal of discussion of recent years
upon the subject of zig-zag leakage. Some authors hold that the only cross
flux of this kind which should be taken into account is the flux which passes from
stator to rotor backwards and forwards, interlinking with some of the stator and
some of the rotor conductors, due to the fact that back magnetomotive force of
the rotor currents is not everywhere balanced by the magnetomotive force of the
stator currents. This cross flux may be spoken of as the " doubly interlinked
leakage." In the opinion of other authors, there is a cross flux which zig-zags
backwards and forwards across the air-gap by reason of the fact that at certain
positions of the rotor teeth with respect to the stator teeth the open slots of the
stator are in a measure short-circuited by the tops of the rotor teeth, and the open
slots of the rotor are in a measure short-circuited by the tops of the stator teeth.
The amount of the short-circuiting is a function of the numbers of teeth on stator
and rotor, of the widths of the tops of the teeth, and the length of the gap. This
true zig-zag leakage would occur however 'well balanced the stator and rotor
magnetomotive forces might be (provided always that the tops of the teeth were
staggered for some part of the time of revolution, as indeed they must be).

We will give a simple rule for the rough estimation of the zig-zag leakage which
works well enough in practice ; though by reason of the fact that it does not take
into account all the factors which aflect the result, it cannot be regarded as strictly
accurate. As we said before, if a method is not short, it is of no use in practical
design. The rule here given sacrifices all the minor refinements in order that it
can be applied in 30 seconds. If the reader requires a more exact method, he
is referred to Dr. Goklschmidt's paper mentioned on page 421.

The reluctance of the path of the zig-zag leakage is in the main proportional
to the length of the air-gap. The width of the path changes as the teeth change
their relative positions ; but the maximum width of the path is one-half the width

424

DYNAMO-ELECTRIC MACHINERY

of the tops of the teeth where these aie equal in rotor and stator, and where these
are unequal it is a function of the widths of the tops of the teeth.

If we assume that the dimensions of the teeth and the mouths of the slots
are such as one generally finds in practice, it is possible, roughly, to take into account
the changing width of the leakage path by means of a coefficient Kz, and we may
write :

\ _.jy pitch of slot 1

* "" * 2 length of gap x Kg

where X^ denotes the lines of zig-zag leakage per cm. axial length of slot, for
unit magnetomotive force applied across the mouth of a stator slot. The values
of Kz which may ordinarily be employed in practice are given in Fig. 403 as a function

r.f 4.1*^ ,«x'^ No. of stator slots
of the ratio -== -. — =— — .

No. of rotor slots

•6

^5 W ii K S 16 20 2*2 2^ 26 2€ 90
' Number ofStatorSLots

Number of Rotor Slots

Fio. 403. — ^Values of Kt for eBtimating dg-zag leakage.

Example 52. The slots in the stator and rotor of a 1500 h.p. motor are shown in Fig. 408.
The air-gap ^^=0*2 om. ; the contraction cbefBcient Kf =1'2; the pitch of the stator slot8=2*6
cms. There are 288 slots in the stator (3 conductors per slot) and 360 slots in the rotor. Find
the zig-zag leakage per pole for a core length of 47 cms. when the motor is on full load of
260 amperes per phase.

OQQ

=0-8, and from Fig. 403 ^,=0-34,

360
X,=0-34x

2-6

2x0-2xl-2

= 1-84.

If we now add together the permeances due to the stator slot, the rotor slot
and the zig-zag path per cm. of axial length, and multiply by twice the length of
iron, we arrive at an approximate figure for the permeance of the path of magnetic
leakage from one pole, so far as the first three parts of the leakage above referred
to are concerned. Leaving out of account for the moment the leakage due to the
end windings, we can get the leakage from the iron paths in c.G.s. lines per pole

INDUCTION MOTORS 426

by multiplying the total penneance above calculated by the maximum ampere-
wires per slot and by 1*257.

Example 53. In the 1500 h.p. motor shown in Fig. 407, we have

Permeance of leakage path aoross stator slot =1*93
„ ,, „ rotor ,, =1-96

,f zig-zag leakage path =1*84

573
Taking the partioulars of the motor given on page 448 :

Axial length of iron =47 cms.

The permeance of the path=5*73x 47x2=540.

For a stator current of 1 ampere the leakage flux along the above paths is

540xlx 1-41 x3x 1-257=2860 lines.

The flux per pole leaking across the iron teeth for one ampere per phase in
the stator we will denote by <^. It is the sum of the slot leakage and the zig-zag
leakage when one ampere is passing in the stator. In the example given above
<^»2860.

Leakage around the end windings. The only really accurate way of finding
the value of the end leakage of an induction motor is by experiment on the winding
in question. If we have two motors built on the same frame with the same type
of winding, but one machine much longer than the other, we can, by measuring
the short-circuit current on each machine, calculate with some accuracy what part
of the leakage reactance in each machine is due to the end windings.

When once this has been ascertained it can be put on record and the figure
used in similar cases.

In default of values found by experiment, it is desirable to have a simple method
of finding roughly the amount of end leakage that may be expected on a given
machine.

It will be seen that, while there are very many types and shapes of windings
on induction motors, there are properties common to all the types found on com-
mercial machines which make it possible to give approximate constants for the
estimation of the end leakage. In the first place, where the coils are very deep
they usually project a very long way out from the core ; so that while the mean
line of path encircling the coils is increased, the area of the path is increased in
about the same proportion. Thus, for a given type of winding, say that illus-
trated in Fig. 201, the leakage per centimetre of perimeter will be about the
same for the same ampere-turns per pole, independently of the size of the coils,
always supposing that they are made to the same drawing, but to different scales.
On the other hand, there is a great deal of difference between the amount of end
leakage &om coils of different tjrpes. It has been found by experiment (as is, indeed,
obvious from inspection) that coils of the barrel type, as illustrated in Fig. 129,
do not give half as much end leakage as coils of the concentric or chain type, as
illustrated in Fig. 114. It will be sufficient for our purpose to introduce certain co-
efficients to take care of the characteristics of the different types of coils, and to
include in our formula only those factors which have the greatest influence on the
leakage per pole, assuming that the coil is of a standard type. As we are concerned
in this formula with the leakage per pole, one of the main factors is the pole pitch.

426 DYNAMO-ELECTRIC MACHINERY

Where the pitch is short and the coils project a long way from the iron, there is
a great deal of sidewaye leakage that ought to be taken into account in the formula.
The amoimt of end leakage depends, not so much upon the number of ampere
wires per slot, as upon the total number of ampere-tums per pole. The nearer the
rotor and stator windings lie together, so as to neutralize each other in the creation
of a magnetic field, the less will be the end leakage. Thus, if we have a barrel
winding on Imth stator and rotor, the end leakage will be much less than if both
rotor and stator windings are turned away from each other towards the iron. The
further the windings project from the &ame, the greater will be the leakage. The
proximity of the iron parts, including the end plates and fenders covering the
winding, greatly affects the end leakage. The whole matter is so complicated by
accidental circumstances that it is useless to attempt any accurate calculation.

FlO. «M.— SbowiBB dimi
Inductk

In order to arrive at some rough idea to serve as a basis of calculation, we may
divide the types of end windings into four separate classes, as shown in Table
XVIII. To each combination of one type of stator winding with one type of rotor
winding we may attach the coef&cient Ki given in the table. These coefficients
can then be used in conjunction with the following formula :

End leakage in c.o.s. lines per pole on both ends of machine,
laitr = Kix (Ip + Op) X virtual a.t. per pole,
where Ki has a value somewhere between TS and 35, depending on the type of
winding, as shown in the accompanying table, and
ip— pitch of poles in cms.,
a,=average overhang of coils in cms.

In Fig. 104 the average overhang of the coils is 100 mm., so that Oi^lO cms.

The virtual a.i. per pole are taken in the following manner: Take the total
number of conductors per phase per pole and multiply by the virtual amperes per
conductor.

INDUCTION MOTORS

427

The end leakage on one pole really depends on how the end windings are arranged
on that pole. There will be a difference, for instance, between the amount of flux
encircling the hemitropic winding shown in Fig. 101 and the flux encircling the
divided -coil winding shown in Fig. 102. If, however, we take the ampere-turns as
directed above, and remember that it is the total leakage on two poles that must
be taken into consideration, it will be found that the above method gives values
which are near enough for practical calculations. The hemitropic winding usually
has a larger a^ than the divided coil winding, and in that respect gives rather
greater values for end leakage.

Table XVIII. Values op Kl for End Leakage or Thbee-Phasr Motors with

Normal Full-pitch Windings.

Typb or Stator Winding.

Ttpk of Rotor.

Barrel
(Fig. 129).

Muxh
(Fig. 138).

Ck)Dcentric
(Fig. 114).

Squirrel cage (Fig. 413) -
Barrel (Fig. 129) -
Mush (Fig. 138)
Concentric (Fig. 114)

1-8
1-4
2-2
2-45

26
2-4
S-1
3-2

2-8
2-45
3-2
3-5

Example 54. In the 1500 ii.p. motor, particulars of which are given on page 448, the pitch
of the poles /^ is 31 cms. and the average overhang a^ of the coils is 12*5 cms. There are 4 slots
per phase per pole, and 3 conductors per slot, so that for 1 ampere per phase we have the
virtual a.t. per pole =4 x 3 x 1 = 12.

The type of winding is "concentric" on the stator and "barrel" on the rotor, and from
Table XVIII. we get Kl=2'45, Therefore the end leakage per pole is

^.=2-45 X (31 + 12-5) X 12= 1275 c.o.a. lines.

We will denote by </>« the end leakage per pole when one ampere per phase is
passing in the stator winding. Then Ia<t>e is the end leakage for any current la.

As we have seen, the short-circuit current of the motor depends mainly upon
the value of the sum of all the leakage fluxes for one ampere passing in the stator.

We will write <^i + </>« = <^(, the total leakage per pole for one ampere in the
stator.

In the above examples 0i-|-^e=4]35, and at no load with 90 amperes per phase we have

9O(0( + ^.) = 9O0 =372x 10»,
the total leakage for 90 amperes in the stator.

Then, if <l>p is the total flux per pole at normal voltage,

^ = /,, the short-circuit current,
91

when normal voltage is applied to the short-circuited motor, assuming that we
can neglect the resistance (see page 428).

Example 55. In the before-mentioned 1500 u.p. motor with 90 amperes per phase in the
stator, the total leakage flux is 3*72 x 10^. The leakage for one ampere is

01=4135.
Now the total flux per pole at 3000 volts, 0^=5*6 x 10*. We have then

0p_5:6xiO»_
0,-^135^""^^ """^^

428 DYNAMO-ELECTRIC MACHINERY

This IB the short-circuit ouzrent there would be if there were no resistance. The ratio of the
magnetizing cuirent to this I, is sometimes denoted by r.

Thus, r=^,

_ leakage flux at no load
total flux per pole

0.70V m>
In above example r =^^^^^ =0 067.

THE REACTANCE OF THE MOTOR ON SHORT CIRCUIT.

Having calculated the current that would flow if there were no resistance, we
can at once get the reactance of the motor regarded as a short-circuited transformer.
We can write :

Volts per phase = short-circuit current x reactance per phase.

Example 56. In the motor described on page 448, the voltage per phase is 1730, and from
the calculation of the leakage flux, the current per phase, if there were no resistance, would
be 1350 amperes. ^^^^ 1360 x a:.,

a:« = l*3.

THE APPARENT RESISTANCE OF THE MOTOR ON SHORT CIRCUIT.

An induction motor with its rotor locked, that is to say, held so that it cannot
turn, and with the rotor circuit closed on itself, behaves like a short-circuited trans-
former. The apparent resistance observed at the terminals of the primary depends
upon the resistances both of the primary and secondary windings. To obtain the
efiect of the rotor resistance, as observed at the terminals of the stator, it is necessary
to multiply the actual resistance of the rotor windings by the square of the ratio

of transformation -^, Let r^ be the resistance per phase of the rotor winding ;

then s 2

^2
as the apparent resistance of the rotor observed at the terminals of the stator.

The total apparent resistance per phase ta is *"! + o'2^2' where r^ is the resist-
ance per phase of the stator (see page 456). ^

THE APPARENT IMPEDANCE OF THE MOTOR ON SHORT CIRCUIT.

Having calculated the apparent reactance of the motor windings per phase
(see page 455), we can obtain the apparent impedance by the formula

Example 57. In the 1500 h.p. motor described on page 448, the number of conductors in
the stator in series is 864, and the number in the rotor 360. Both are star connected,

therefore 7t=k5r=2*4. The resistance of one phase of the stator winding is 0*074 ohm, and
02 »>ou

the resistance of the phase of the rotor winding is 0*0133. Therefore

r2.i = (2-4)«x0-0133=0'0766.

INDUCTION MOTORS 429

Thus f ^ = f 1 + fj. 1 = 0-074 + 00766 = 0-1506.
And we have found that a;a^l*3.
Therefore the apparent impedance

r=V0023 + 1-69 = 1-31

The short-circuit current can be obtained by dividing the voltage per phase by
the apparent impedance per phase.

Example 58. In the 1500 h.p. motor to whioh the above examples refer, the terminal

voltage is 3000» and as the stator is star-oonneoted the voltage per phase is 1730. The short-

E 1730
oirouit current /•a=^=Y:qT -^^20 amperes per phase.

Having calculated the no-load current and the short-circuit current, the Heyland
diagram can be constructed as described on page 414, and from it we can obtain
the power factor, the slip and the e£G[ciency at various loads — ^the starting current,
the starting torque, the maximum torque and the maximum output.

An example will be found fully worked out in connection with the 1500 H.P.
motor described below.

The power fGu^tor for various values of r and various loads. It will be seen
that if the ratio between the no-load current In and the short-circuit current /«.
Ib fixed, then the power factor for a load forming any specified fraction of the maxi-
mum load can be determined from the Heyland diagrams. The only other quantities
which would affect the power factor are the resistances of the stator and rotor,
and if these are small they affect the result to a very small extent.

In order to be able to state what the power factor of any motor will be at any
particular load, without going through the calculation, it is convenient to have
curves such as those given in Fig. 405. Each curve is drawn for a different value
of T, where

_ leakage flux at no load
~ total flux per pole.

magnetizing current

~ wattless component of short-circuit current'

As the resistance of the stator and rotor windings does have some effect upon
the power factor, the curves have been drawn on the assumption that the power
factor on short circuit is 0*25 ; that is to say,

r^=0-25xa.

This is not so far from the truth in many commercial motors as to call for any
correction of the curves where the power factor is less than 0'25 on short circuit.
Where, however, the power factor on short circuit is as great as 0-5, we should use
the curves as if the values of t attached to each curve were increased 10 %.
Thus, power factor =0*5 on short circuit, and the top curve should be used for
T=0022.

The crawling of induction motors. When a squirrel-cage motor is being started
up it sometimes attains about one-seventh of full speed and refuses to go any

'^.

Load as a Fiuction, vf Maximum Ouiput; fI^pawerfycu>r-o-zsJ

FIO. 105. — CiitT« giving tbt power Iictor ol u Inauctloii molot

«5.

Load eis a Frttction of Maximum. OutpiUfd^pmerfyOor-o-zs/

whSD loulBd with uiy given InctloD ol

432

DYNAMO-ELECTRIC MACHINERY

faster, until by some means the speed is carried over the dead point, when the
torque of the motor increases and it runs up to fall speed. This .is due to the
presence of a pronounced seventh harmonic in the field-form (see page 22). The
harmonic has the effect of superimposing on the main field, and the field having
seven poles within the span of one main pole pitch. As these poles alternate
with the frequency of the supply (say 50 cycles), they produce a torque on
the rotor having the same characteristics as the torque produced by a motor
having seven times as many poles, and having a sjmchronous speed one-seventh
of the normal speed. The whole torque on the rotor is thus made up of
the torque, due to the main revolving field plus the torque due to the seventh
harmonic.

Fig. 406 shows the speed-torque curve of such a motor. The torque is greatest
when running slightly under synchronous speed. If we drive the motor above

5CAbU

OACKWAMS

200%

Fio. 406. — Showing how the torque-speed chazactertotic of a BqnJrrel-eage motor to affected

by harmonics in the field-form.

synchronous speed, the torque becomes negative. Now the torque due to the seventh
harmonic is similar in shape, as is shown by the dotted line which crosses the zero
line at one-seventh of full speed. If this dotted curve be superimposed upon the
main characteristic, we get the curve shown by the full line. We see, therefore,
that as the motor starts from rest the torque increases up to the little peak on the
curve, and then rapidly diminishes as we approach what is the synchronous speed
for the seventh harmonic. Above that speed the torque still diminishes, because,
so far as the harmonic is concerned, we are getting a braking action. It is not
until we have increased the speed by an amount which takes it well past the seventh
speed that the torque again begins to increase. If, now, the friction of the motor
should be such that it requires a torque greater than that supplied by the motor
in the region of the seventh speed, the motor will crawl round and be perfectly
stable in maintaining its speed between certain limits. If we have a fifth harmonic

INDUCTIOl^ MOTORS 433

on the field-form, it tends to produce rotation in the opposite direction, and
gives to the motor the characteristic shown in Fig. 406. These matters are very
lucidly discussed by Mr. Catterson Smith in a paper* from which Fig. 406
is taken.

In order to avoid this crawling of squirrel-cage motors it is necessary to make
the resistance of the end rings of the cage- great enough to give to the motor at
seventh speed a torque well above the friction torque.

Slip of an induction motor. Although the slip is given by the ratio oi XY : PY
in Fig. 400, it is not convenient to calculate the slip by scaling off these vectors,
because XY is too small to be accurately measured. A much more accurate way is
to calculate the ohmic losses in the rotor from the known resistance and the known
current. For this purpose we may either take the current by scaling off PN (in
which case the ohmic loss will be SPN^ r^^i), or we may multiply the current PN by
the ratio of transformation SJSiy and obtain the true rotor current I2, (in which
case the ohmic loss will be ^I^r^, In this we have taken r^ as the resistance of one
leg of the star-connected rotor winding. Having obtained the ohmic loss in the
rotor 3/^/2, we obtain the slip from the formula ;

Output in watts-hS/jfj
Examples are given on pages 459 and 467.

* Catterson Smith, Journal Inst, Electrical Engineers, vol. 49, p. 635, 1912.

" • B(.

CHAPTER XVII.

THE SPECIFICATION OF INDUCTION MOTORS.

The uses to which induction motors are put are so very various, and the performance
required in the various cases so very difEerent, that it is difficult to find a wholly
satisfactory classification when we come to treat of the subject in a systematic
way. Broadly, there are two classes : (1) motors that are started and stopped only
once or twice a day, and run at an almost constant speed ; and (2) motors that
are started and stopped or reversed frequently, and may be required to run at
various speeds. The first class we find driving machines that are running all the
time on steady or varying loads, such as motor generators, pumps, the counter-
shafts in small factories, the main shafts in large mills that have been converted
to electrical driving, cotton-spinning nxachinery and flour mills. The second class
we need for machine tools that are often started and stopped, cranes, lifts and
winding engines. An intermediate class, so far as service is concerned, is formed
by the motors which, while running all day in one direction, have their speed changed
over a wide range, as, for instance, non-reversing rolling-mill motors, which must
permit a flywheel to give up a large part of its energy.

Small motors for the first class of service are often made of the squirrel-cage
type with low-resistance rotors and high efficiency. They are usually started up
without much load, either by means of an auto-starter or by connecting the windings
of the stator first in star and afterwards in mesh. Such motors start up on about
full-load current when provided with an auto-starter, and on less than three times
fidl-load current with the star-mesh system. Where a small motor has to be started
up on load the squirrel cage can still be used, if the amount of current drawn horn
the line does uot matter, or if the efficiency does not matter. In the latter case
the squirrel cage is made of fairly high resistance, so as to give a good starting
torque. Large motors for constant speed, if they can be started light, are some-
times made of the squirrel-cage type. But the advantage of this type in the case
of very large motors is doubtful. The mechanical construction of the squirrel
cage to deal with very large currents, and provide properly for expansion and
contraction, is hardly any easier or cheaper than a barrel winding, and the starting
with a wound rotor is so much more satisfactory that most large motors are now
made of that type. Even when the motor is started up independently (as in a
motor-generator started on the continuous-current side) and switched on at full

THE SPECIFICATION OF INDUCTION MOTORS 435

speed, we do not get rid of shocks at the instant of switching unless we employ
a charging coil or water resistance.

For crane work and for motors for hard service in dirty situations, where the
efficiency is not of first importance, squirrel-cage motors with high-resistance rotors
are sometimes employed, but the methods now used of enclosing motors and slip-
rings are so efficient, and the construction of wound rotors so hardy, that there is
now a great deal to say in favour of external resistances instead of the high-resistance
squirrel cages.

For the second class of service, rotors with a wire or bar winding, and provided
with slip rings, will be' used to enable the resistance of the rotor to be increased
at starting and when running much below s3mchronous speed. Motors of this
type can be started up on full-load torque with not much more than full-load current,
and at nearly full-load power factor. Where desired, the starting torque can be
made three or four times the full-load torque, with a corresponding increase in the
starting current. To call for more than three times full-load torque usually involves
pa3ang the price of a motor built on a lai^er frame, and there will be a consequent
loss in efficiency.

General. The specification, besides stating the voltage and frequency of supply,
should give such particulars of the source of power and the other apparatus in
circuit as may be necessary to enable a manufacturer to judge whether his motors
are suitable for the circuit in question. If the motor is to be run from town mains,
there will generally be some restriction as to the current that may be drawn and
its power factor. If there is a long feeder in circuit likely to cause a serious drop in
the voltage on load, the fact should be stated. Particulars should be given of the
general character of the load and the probable accuracy with which the horse-
power and maximum torque have been estimated.

Starting. The method proposed for starting the motor should be stated, together
with the starting torque required and the amount of current that may be drawn
from the line. The power factor of the motor at starting is a very important
characteristic affecting the cost of the installation. If the motors are small, as com-
pared with the whole power of the circuit to which they are connected, and there
are no special circumstances which call for a good power factor on starting, it is
not well to insist on too stringent conditions, or the complication and cost of the
plant may be unduly increased. It is, for instance, quite a common practice to
start up 30 and 40 h.p. squirrel cage motors in a large factory on simple auto-
starters or on the star-mesh method, and though the rush of current at low power
factor is considerable, it does not affect the good working of the whole factory,
and the plant is simpler than if slip-ring motors had been employed. The maximum
torque required should be stated with as great accuracy as possible. It must on
no accoimt be understated, because the maximum torque which an induction motor
can give is a fairly definite quantity, and cannot be increased by merely over-
loading the motor as with a direct-current motor. On the other hand, it should
not be overstated, or the manufacturer will supply a motor which is too large for
the work, so that the efficiency and power factor will not be as good as they might be.

Speed. If there are any reasons why the speed must be kept exceedingly constant,
the fact should be stated. Usually a statement of the purpose for which the motor

436 DYNAMO-ELECTRIC MACHINERY

is requiied is sufficient information for the manufacturer to go upon in adjusting
the slip of the motor, and it is only in special cases that it is necessary to specify
the slip exactly.

Where wide variations in speed are required, the range of speed should be given
as accurately as possible, and a statement should be made whether the change of
speed must be continuous over the whole range, or whether it is permissible to
introduce pole-changing, gear-changing, or any devices which will increase the
efficiency at certain speeds, or in any way reduce the difficulty of obtaining the
wide range of speed.

Power fEbctor. It is usual to permit the manufacturer to specify the power
factor of his motor. It is sufficient for the purchaser to indicate the relative im-
portance of a good power factor in his particular case. The manufacturer can
then choose a standard motor to meet the case, without being compelled to alter
the windings to aim at some particular figure. It should not be forgotten that
where a good power factor is of the greatest importance, devices can be added
to the motor which will make the power factor unity or even leading (see page 605).

MaTimum torque. The maximum running torque which a motor is to yield is
a very important consideration in determining the size of the frame upon which
it must be built, and therefore the price of the motor will largely depend upon the
maximum torque required. For general work a torque of 2^ times the running
torque is considered sufficient to prevent accidental pull-outs, but the circumstances
of each case should be considered, and where severe over-loads are likely to come
on, the motor must be designed to meet them. In some cases the load may be so
steady that a much smaller maximum torque may be sufficient, and a saving can
be made both in the cost of the motor and in the power taken to drive it.

Temperature rise. What was said on page 256 about temperature rise is also
applicable to induction motors. With crane motors of the squirrel-cage type,
provided with high-resistance end rings, it is usual to allow the temperature to rise
to 150'' to 200" G.y the construction being specially designed to withstand these
temperatures without injury.

Puncture test. The rules with regard to puncture test are in general the same
for the stator of an induction motor as for the armature of a generator. In cases
where it is intended to switch an idle motor directly on to a high-voltage line,
special care is necessary in the insulation between turns of the coils of the stator
nearest the terminals. In such cases it may be necessary to specify a certain test
between turns on these coils, the test to be made during the course of construction.
If the insulation is designed to resist an instantaneous puncture test between turns
of one-half the voltage of the motor, it will in general be sufficient to withstand
being switched suddenly on to the line.

Arrangement of f^ame and shaft. The specification should state any matters
relating to the arrangement of the frame which are important for the installing
of the motor. It should state, for instance, whether the motor must go on its own
bedplate, or whether the frame must be designed to fit some special support.

The bringing out of the terminals is a matter which in some cases requires
special consideration, particularly with motors that are to be put in rather inacces-
sible places. The specification should draw attention to any points of this kind.

THE SPECIFICATION OF INDUCTION MOTORS 437

Then, again, it sometimes happens that a motor requires special protection from
dirt or dripping water. Protection on one side may be sufficient, or a total enclosed
motor may be required.

The specification should also give particulars as to how the motor is to be
connected to the load, whether by pulley and belt or spur gear, and of the sizes
of these. If the motor is to be direct coupled, particulars should be given of the
kind of coupling and its size, and a statement made as to how much of the coupling
is to be supplied by the manufacturer of the motor. Sometimes it is necessary
to have the shaft of extra length, or turned to a special size or shape. Matters of
this kind should always appear on the specification, as they have a very consider-
able efiect upon the cost of manufacture.

438

DYNAMO-ELECTRIC MACHINERY

SPECIFICATION No. 7.

Extent of
Work.

Fnnction of
Motor.

Type of
Botor.

Characteristics
of Motor.

1500 H.P. INDUCTION MOTOR

85. This specification covers the manufacture, supply,
deUvery, erection, testing, and setting to work of a three-
phase induction motor, direct connected to a 1000 k.w. con-
tinuous-current generator in the Sub-station of the
Corporation in Street,

86. The motor is intended to drive an existing continuous-
current generator of 1000 k.w. capacity at a speed of 246
revolutions per minute. The said generator feeds the mains
of the power and lighting supply of the town of with
continuous current at a pressure of 500 volts.

87. The rotor shall be of the wound type provided with
slip-rings.

88. * The motor is to have the following characteristics :

Normal output 1500 h.p.

Normal voltage at ter-
minals

Frequency

Number of phases

Speed

Power factor not less
than *

How connected to load Direct connected through flange

coupling.

Temperature rise after

6 hours full load nm 40"^ C. by thermometer.

3000 volts.

60 cycles.

246 revs, per minute.

Over load
Temperature rise after

3 hours 25 per cent.

over load
Maximimi torque
Starting torque

25 per cent, for 3 hours.

55° C. by thermometer.
2^ times full-load torque.
Sufficient to start the 1000 K.w.
generator unloaded.

* The Contractor is to state the power factor at full load of the motor he proposes
to supply.

THE SPECIFICATION OP INDUCTION MOTORS 439

Puncture test 6600 volts alternating at 50

cycles applied for 1 minute
between the stator windings
and frame.
3000 volts alternating at 50
cycles for 1 minute between
rotor windings and frame.

89. The contract includes the delivery of the motor at the ^^;^^^^
Sub-station of the Corporation, together with bedplate, bear-
ings and pedestals, and the erection, aligning and coupling

of the same to the 1000 K.w. generator. The switch gear and
starting gear are provided for under another specification.

90. The stator frame shall be of the best cast-iron, of stator
deep section and great stiffness, so as to prevent any appreci-
able distortion due to magnetic pull. The finger-plates
supporting the ends of the stator and rotor teeth shall be of
very rigid construction, and shall be approved by the Pur-
chaser.

91. The parts of the stator coils projecting from the stator cofls.
slots shall be so rigid that no appreciable movement of them
occurs under the most severe conditions of service.

91a. Each coil shall be wound so that those conductors Arrangement
between which the highest potential occurs are furthest "" *"" "^^"*
separated from one another.

92. The stator shall be fitted with strong guards or Fenders,
fenders to prevent the h.t. windings being accidentally
touched by hand. These fenders shall be designed so that
they do not interfere with the ventilation.

93. The individual conductors forming each stator coil insulation oi
shaU be insulated with mica bound in position in an approved

way. The coils shaU. be dried and impregnated under vacuum
with insulating compound ; they shall then be wrapped on
the straight portions with mica mounted on cloth or paper,
the whole moulded under pressure so as to exclude air-spaces.
The ends of the coils shall be insulated with tape treated
with a suitable insulating varnish of the highest quahty.
The stator slots shall be lined with paraffined fullerboard to
prevent abrasion of the tape when the coils are inserted.
The whole of the insulation shall be carried out so as to be

i40 DYNAMO-ELECTRIC MACHINERY

permanent and reliable, and so as to withstand well the
heating and vibration to which the coils may be subjected.
After being placed in the slots, the coils shall be completely
covered with a non-hygroscopic waterproof varnish capable
of withstanding the action of hot oil, and having a smooth
surface which will neither soften nor crack under working
conditions.

pre«iire Tenti'. 94. All prcssuTe tests hereinafter specified shall be carried
out with an alternating voltage at a frequency of 50 cycles
per second.

T^ on 95. Before being placed in the slots, each stator coil shall

winding. be tested at 2000 volts between each pair of adjacent turns.
After the coils have been placed in the slots and before they
are connected up, the whole of the coils shall be subjected
for 1 minute to 8000 volts between copper and iron. After
the coils have been connected up, but before the phases are
interconnected, they shall be tested between phases A and
B, B and C, and C and ^ at a pressure of 8000 volts. After
the motor has been delivered and run on full load for 6
hours, it shall, while still warm, be subjected to a pressure
of 6600 volts, for 1 minute, between stator copper and iron.

Botor. 96. The rotor shall be built upon a cast-iron spider with

free arms, designed to avoid excessive stresses arising through
the cooling of the casting. The rotor winding shall consist
of copper bars which are insulated before being placed in the
slots ; each bar shall be insulated on the straight portion by
mica mounted on paper or cloth, and held in position by tape.
The end portions shall be insulated with Empire cloth and
treated cotton tape. The insulation of the bars shall be com-
pletely impregnated with varnish and well dried out. The
slots shall be lined with parafiined fuUerboard to prevent
abrasion of the tape when the coils are inserted.

Botor ci)Us ^^' ^^^ ^^^ rotor bars have been connected together,

but before they are connected in star or in mesh, a pressure
of 4000 volts shall be applied for 1 minute between phases
A and B, B and (7, and C and A. The phases shall then be
connected in star or in mesh, and a pressure of 3000 volts
shall be applied between copper and iron. This test shall
be repeated after the motor has been run at full load for
6 hours and while it is still hot.

THE SPECIFICATION OP INDUCTION MOTORS 441

98. The tender shall state the radial clearance between the Air-gap.
stator and rotor iron.

99. The efficiency of the motor shall be calculated from the Efficiency.
separate losses, which shall be measured in the following way :

1. Iron loss ^ friction and windage. The machine shall be
nm unloaded at 3000 volts between terminals, and the power
taken to drive it measured by the two-wattmeter method.
The sum of the readings shall be taken to be the iron loss,
friction and windage.

2. Copper losses. The rotor shall be locked and the rotor
windings short-circuited through suitable ampere-meters. A
voltage shall then be applied to the stator at 25 cycles * per
second and gradually brought up until the rotor winding
yields full-load current on short circuit. Measurement shall
be made of the power supplied to the stator under these con-
ditions by means of two-wattmeter readings ; the sum of
these readings shall be taken to represent the copper losses
at full load.

100. The Contractor shall guarantee the efficiency calcu- Guarantee of
lated from these separate losses, and he shall further guarantee ^"^^^"^^
that when the motor is in operation the over-all efficiency
actually obtained shall not be more than 1 per cent, lower

than the efficiency so calculated. The figures for the calcu-
lated efficiency shall be given at full load, three-quarter load
and half load.

101. In addition to the puncture tests specified in Clauses Tests at
95 and 97, the following tests shall be made at the maker's
works :

1. Iron loss, friction and windage test as specified in
Clause 99.

2. Copper loss test as specified in Clause 99.

3. Resistance test. The resistances of the rotor and stator
windings shall be measured.

102. The following tests shall be carried out after the Tests on site.
motor is erected in the Sub-station of the Corporation :

1. Temperature test. The motor shall be run for 6 hours
at full load, which for this purpose shall be taken to mean

* The reason for specifying that this test shall be carried out at one-half the rated
frequency is that the losses on the rotor at a high frequency' would be unduly
increased.

442 DYNAMO-ELECTRIC MACHINERY

an input of 1200 K.w. at 3000 volts. At the end of this run
the motor shall be stopped and temperatures taken with all
possible speed. The temperature of any part of the motor,
as measured by a thermometer, shall not rise more than 40° C.
above that of the surrounding air. For this purpose the
temperature of the air shall be taken to be the temperature
measured in Une with the shaft of the motor at a distance
of 3 feet from the end of the shaft.

2. Over-load test Inunediately after taking the tempera-
tures, the motor shall be put for 3 hours on over load, which
for this purpose shall be taken to be an input of 1500 k.w.
at 3000 volts ; after which run the temperatures shall be
taken again. The highest temperature rise shall not exceed
55° C.

3. Power-factor test. During the temperature run the
power factor of the motor shall be measured by means of a
power-factor meter, and also by the two-wattmeter method.

starting. 103. The motor will be started with a starting resistance

described in another specification, connected in series with
the rotor.

Power Factor at 104. Thc ciuTent drawn from the line shall not at any
staring. ^^^^ duriug thc Starting exceed full-load current, and the
wattless current shall not exceed half full-load current.

Slip Rings. 105. The sUp rings on the rotor shall be of substantial

design, and shall have a wearing depth of not less than 1^ in.

Brush Gear. 106. The brackcts supporting the brush gear shall be of

very rigid construction. The brush holders shall constrain
the brushes so that they always slide in a direction parallel
to the same line. The brushes shall be of the metal-carbon
tyr® by a good maker, and they shall be fitted with a flexible
connection capable of carrying the current on 25 per cent,
overload without undue heating. There shall be absolutely no
sparking on the slip rings when the machine is running on
25 per cent, overload.

Short- 107. The brush gear and slip rings shall be provided with

Device"* a dcvicc whcrcby the slip ring may be short circuited when the

motor is up to speed, without the current passing through the

brushes.

THE SPECIFICATION OP INDUCTION MOTORS 443

108. The design of the fenders around the stater winding ventnation.
and the shaping of other parts of the motor shall be such that

the air thrown out by the rotor shall be thrown out well to
the surrounding atmosphere, and shall not to any appreciable
extent be thrown toward the continuous-current motor, or be
caused to circulate in an eddy so as to return on the motor
itself. While ample ventilating ducts must be provided in
stator and rotor, these must be designed so as not to create
excessive noise.

(See Clauses 6, p. 271 ; 36, p. 360 ; 74, p. 382 ; 272, p. 591.) Foundations.

(See Clauses 55 to 59, pace 379.) Accessibility

^ » r © / of Site.

(See Clauses 8, p. 271 ; 60, p. 379 ; 273, p. 591.) Use of Crane.

109. The rotor shall be well balanced, and when at full Balance,
speed shall not communicate to the bearing pedestals any
appreciable vibration.

110. The shaft of the rotor shall be fitted with a half coup- coupling.
ling (forged with it) for coupling to the shaft of the con-
tinuous-current generator. The coupling bolts and nuts shall

be completely covered by a steel shrouding. All coupling
holes shall be reamered out in position, and well-fitting
finished bolts and nuts shall be supplied. After fitting, each
coupling shall be clearly marked for correct matching when
re-erecting.

(See Clauses 67, p. 380; 268, p. 590.) Bearings.

111. The motor shall be mounted on a bedplate and two Bedplate,
pedestal bearings. The bedplate shall be arranged to be
bolted to the existing bedplate of the continuous-current
generator, particulars of which are given in drawing No.

The height of the pedestal shall be arranged so as to bring the
rotor shaft in line with the existing shaft of the continuous-
current generator.

(See Clause 186, page 523.) Hoiding^iown

Bolts.

112. The position of the terminals shall be shown on the xenmnais.
tender drawings. The connections to the high-tension cables

shall be made by means of sweated thimbles, which shall be
completely insulated by means of substantial insulating
sleeves. The cable which makes connection between the

444

Sparen.

DYNAMO-ELECTRIC MACHINERY

stator winding and the terminals shall be insulated with
Empire cloth or other insulation which does not soften with
heat.

113. The terminals of the stator winding shall be brought
by means of cables insulated by Empire cloth, or other
insulation which does not soften with heat, to a cast-iron
terminal box, in which the terminals shall be mounted on
independent porcelain insulators. Wide and efficient insulat-
ing screens shall be placed between the terminals of the three
phases, so that arcing between phases is impossible. The
connections from the high-tension switchboard will be made
by means of a three-core paper-insulated cable, which will
be brought to a trifiircating box designed to fit on to the
terminal box aforesaid, for convenient connection between
the high-tension cable and the terminals of the stator.

114. The Contractor shall supply the spare parts set out
in Schedule I.

Tools.

Maintenance
Period.

Cleaninff and
Palntinic

Drawinffs
supplied with
Hpeciflcation.

116. The Contractor is to provide a full outfit of the
spanners and special tools necessary for disassembUng and
assemblmg the motor, together with a rack for holding them.

(See Clauses 1386, p. 469 ; 206, p. 528.)
(See Clause 209, page 528.)

116. Drawing No. suppUed with this specification
shows the existing lay-out in the Sub-station of the Corporation
and the proposed site for the induction motor.

117. Drawing No. gives particulars of the existing
contiauous-current machine, with bedplate and outboard
bearing, to which it is proposed to connect the induction
motor.

(Contractor to 118. Thc Coutractor is advised to inspect the site and make

Mca8°uroinonti. all uecessary measurements. The Contractor is to be respon-
sible for obtaining any information which shall be necessary for
him in deciding as to the suitability of the site for his plant,
and also for the exact dimensions of all bedplates, bearings,
heights, clearances and foundations, and other matters with
which he may be concerned.

' THE SPECIPICATION OF INDUCTION MOTORS 446

119. Schedule No. II. gives a list of the drawings and i>r»wtop to be

suddIicu wiui

samples which are to be submitted with the tender. Tender.

(See Clause 212, page 529.) Provisional

SCHEDULE No. I,
List of Spares.

SCHEDULE No. IL

List op Drawings required with Tender.

1. Outline drawings showing in plan and elevation the
dimensions of the proposed motor coupled to the existing
CO. generator.

List op Samples required with Tender.

1. Sample of stator coil, showing method of insulating
between copper and iron and the method of insulating the
bent part of the coil. The coil shall also show the method of
making joints in the conductors.

2. Sample of rotor bar with its insulation.

3. Sample of brush holder and brush for sUp rings.

design of 1500 H.P. INDUCTION MOTOR

3000 volts, 50 cycles, 246 B.P.M., power factor at full load 0*88. At an efficiency
of 96 % this gives 1350 k.v,a.

We will suppose that it is required to design a motor to comply with the par-
ticulars given in the specification No. 7 (page 438). The first step is to fix upon a
suitable size of frame. In practice, a manufacturer would probably use a frame
for which drawings were already in existence, and upon which he had built similar
motors before ; but if he had no such frame he would take as a guide a suitable
output coefficient upon some such considerations as the following.

The output coefficient of induction motors depends upon a variety of con-
siderations, some of which are the following : the rated output, — the ratio of the
maximum output to the rated output, — the number of poles, — the temperature
rise,— the facilities for ventilation,— the cost, — ^the efficiency ,— and the power factor at
full load. As the interaction of these factors is extremely complicated, it is impossible

446 DYNAMO-ELECTRIC MACHINERY

tq give any concise rules for arriving at the output coefficient in any particular
case. If we took a frame of certain size and wished to get the TnaTimnm pull-out
torque without regard to anything else, we would build an " iron " machine ; that
is to say, we would make the slots very small and shallow to leave room for the
working flux, which would be made as great as possible. The leakage flux being
very small, the diameter of the semi-circle (Fig. 410) and the maximum output
would be very great. As, however, the room for copper would be very small, the
output at which the motor would work in continuous service would be small. To
get a greater continuous output we would increase the size of the slots. This
would restrict the amount of the working flux and increase the leakage, so that
the maximum output would be reduced, while the rated output would be increased.
This would go on until we arrived at a size of slot such as is commonly found in
practice. If the size of slot were still increased, we would have to cut down the
working flux by a percentage higher than the percentage increase in the current
loading, so that the normal output of the motor would be decreased. With the
ratio of current loading to magnetic loading commonly found in 50-cycle motors
of normal speed, the ratio of the maximum output to rated output is about 2 to
2-5. This ratio is found to be quite satisfactory for the ordinary purposes for
which motors are employed, and it gives a fairly economical arrangement of copper
and iron. For very small motors (from ^ to 1^ h.p.) the economical ratio is still
smaller.

For a given frame and given frequency an increase in the number of poles
affects the output coefficient in two ways. The speed being lower, the ventilation
is not so good, and this tends to reduce the output ; but, on the other hand, it is
found that, the pitch of the poles being shorter, the numbet of coils huddled together
is fewer, and this more than compensates for the slower speed, particularly where
fans are added. A comparison of the output coefficients of modem well- ventilated
motors will show that upon the whole an increase in the number of poles increases

the ratio, ^^ P^ , instead of decreasing it, as "might at first be supposed from the

poorer draught of air. This wiU be seen from Table XIX.

The temperature rise guaranteed, of course, affects the size of frame that must
be employed. In what follows we will assume that the guaranteed rise is 40° C.
by thermometer. The flBusilities for ventilation differ very widely with the purposes
for which the motor is intended. Pipe-ventilated motors, for instance, are greatly
dependent on the size of the pipe carrying the air and the pressure used to drive
it. In what follows we will assume that the motor is ventilated with its own fan,
and that there is a plentiful supply of cool air.

Considerations of cost affect the output coefficient, because it does not by any
means pay to make the smallest possible motor to meet the guarantees. Iron is
cheaper than copper, and it will pay better to make a rather bigger " iron " motor
than the smallest possible " copper " motor.

Similarly, the efficiency and power factor required will often compel us to use
a frame of larger size than might otherwise be necessary.

If we take modern^ well-ventilated, standard, 60-cycle motors, running at
speeds that are ordinary for the output, we will find that the ratings of the frames

INDUCTION MOTORS

447

are such as to lead to the coefficients K^ given in Table XIX. The increase in the
number of poles on these standard machines will have the efEect of reducing the
maximum torque, so that while we find that the coefficient Kq is increased slightly
as the number of poles is increased, the maximum load of the frame is at the same
time decreased. If it should be necessary to build a slow-speed motor and
still preserve the large pull-out torque, it wiU be necessary to take a frame of
larger diameter, so as to get more room for increasing the number of slots and the
working of flux per pole. This will have the effect of reducing Kq, In Table XIX.
the coefficient is based on heating considerations as found on standard motors.

Table XIX. Ratikos or Frames of 50-ctclb, 3-phase Induction Motors.

KQxD^xlm X R.P.M. =K.v.A. input.

2>m= diameter of rotor in metres ; /= axial length of iron in metres ; A'o is a coefficient *

applicable under the circumstances set out above.

4 Poles.

6 Poles.

8 Poles.

1
12 Polos. ;

16 Poles.

24 Poles.

K.V.A.

J^o

0-4

1-6

100

1-8

1-85

0-56

0-67

0-6 ;

100

1-7

200

1-96

0-76

0*85

0-9

200

1-8

500

205

116

600

1-9

1000

216

2-2

115

1-26

1-36

1000

1-95

1600

216

2-2

1-3

1-45

1-65

1600

2000

216

2-2

The continuous output of a frame can be increased beyond ratings arrived at
by the coefficient given, by departing from the supposed conditions. For instance,
if we have a motor with few poles which, with ordinary design, gives us a much
greater maximum load than is required for the purpose in hand, it may be possible
on that motor to deepen and widen the slots, and to increase the continuous output
at the cost of the maximum output. An example of this will be found in the
350 H.p. motor, particulars of which are given on page 463. As this motor is not
required to give more than 1-5 times full load, sufficient copper is put into the
stator and rotor to enable it to nm continuously on full load at a rating considerably
higher than would be chosen ordinarily for a frame of this size.

Table XIX . is given for 50-cycle motors. For 25 cycles a motor for the same speed
will have half the number of poles, and the reduction in the number of poles reduces
jBlq, if the sizes of the slots remain as before. On the 25-cycle motor the pole pitch
will be twice as great as for a 50-cycle machine ; and it will generally be good practice
to make the slots considerably deeper than we should on a 50-cycle motor, because
we can do so without making the leakage excessive. It is thus possible to get a
larger number of ampere-wires per cm. of periphery on a 25-cycle motor of the same
output, speed and power factor, than on a 50-cycle motor.

* Note that /)«, and Im are in metres, while D and I given on the calculation sheets are in
centimetres. To arrive at the DH constant, as given on the calculation sheets, we must multiply
the reciprocal of K^ by 10*. Thus, for ^"0=2, we have

^"< "•''•»•• =^=600.000.

K.V.A.

448

DYNAMO-ELECTRIC MACHINERY

Dat*..i6.A7«K^.i9A?.. Type .. €HBflk... ./ISYN MOTOR -IIOTAR¥ .. . . .^T.. .Poles . . -rr»^.Etoc. Spec ..J..

K VJL/^^P... ; P F.rSft Phwe 9 ; Volt* .'9<'<?^. ^ ; Anps per ter /?^/?,...; Octei..'*?^^.....; R P.M.^'f ^.. ; Rotor Amps ^^^.. ;

H-P./^PP-Anps p. cood. ^3/ Amps p. br. arm. — Temp. nae-.^QZC Reculstion ^.^^..OTtTlo$A^?^.%...'^.A<¥!^^

Customer % : Order No ..

.; Qttot. No r...; Perf. Spec.

FIj-wbeet eflTect.

5^?^«™°^^^^^«^i^*~^^^^SS

; poss. laZ«,..-rr— . .

; Circnm. SOO.

K.V.A.

^'9 iff OS

K« :4/ ',....3000. voits=ri*/. x4.!/:€..jc<?.$4.^;?:<3*

Ann. A.T. p. pole.

Max. Fid. ATw.

Armature. "Rev-

o
o

Dia. Outs

Dta. Ins

Cross Length _
Air Vents Z^L

0:e_ ...

..Mean

Opening Min

Aif« Velocity

Net Length 4/ -4 X 89

Depth b. Slots

Section -^SJS _ Vol.
Flux Density.

24. 300

Loss:^^. cu.CV^ Total
Buried Cu.Z5^e-TotaI
Gap ATtn3^,2QO_- wts
VentAreaiML^^i'.;Wts ^ ^^ ^^

Outs. Ares^ajX?^., Wts L -^^<?P^ 1^41 7<?<?

4.7
^: € '

'37 m

__9

_ Q^50
9000
fGJQOO
12.000
4700

I-

NoofSeg!s
No of Sloes
K.

_/^_|Mn.CJrc.
Z3&\x/5 =

Section Teeth .
Volume Teeth-
Flux Density,

Loss'i^p.cu gflLTotal

c
o
o

Weight of Iron —

-Throw

Stir<

Cond. p. Slot

Total Conds

Size of Cond. j3fi. x ./L.

Amp. p. sq

Xength in Slots- 47 —

Length outside ^ Sum

Total Length . _,

Wt. of i.ooo_M<t?_Total
Res. p. i.ooo '^^Total

Watts p._^?Z^rie

Surface p. .(ESlce.

Watts p. Sq. C/77.

Stat.

266
239

p^r.sec

760

432

/s./oo
52.000

~/6,9bo
7600 \

.^fOOjrgs,
M6, 2'/S, 3-/4-

.Q64uZ

.354^ .^

\f93

054)i-2e

H2
f930m

640 fd/Qg,
18Q ■*-4V3=«?74

1040 dq cm.
'PM.\

•00/2

^^e Slots

</v5*

^T*.

r-X-

I
I

', I

* ^

^ "

K--2oa"ii ' ,

; r

4v2

. ^.SffSlots

2^hi6*5700>i-796^i070

rtold Cta% Of Rotor.

Dia.

\ Total Air Gap

Gap Co-eflf. K.

Pole PitchJA? Pole Arc

Flux per Pole .

Leakage n.L . f.l

Area Flux density __

Unbalanced Pull

2M:e

lie

•66

^$^fo\

No. of Seg.L /^„ Mn.Circ.
No.of Slots 360 X '96=

Vents. Z -dC^^

K, ... . ..Section— ._. .

Weight of Iron

.Z2i.

345

390

H4-0Q

jSSO A;gs

A.T. p Pole n.Ixxid
A.T. p. Pole f. Load
Surface . . - ._
Surface p. Walt.-

I'.R

I. R. _..

Amps.

Whunt.

520

No. of Turns ^^ -

Mean 1. Xwm CO/ip. 95CfhS

Total Length _ ._

Resistance ,

Res. per i .000 | * 23^

Size of Cond-

Conds. per Slot

Total

Length

Wt. per 1,000.

5x/-5

2
720

^iSS

73s\Kcm.

_-*

Total Wt. ^

Watts per Sq

Star or Mesh

Paths in parallel

630
4-30
•066

Magnetization Curve.

Core

Stator Teeth
Rotor Teeth
Gap

Pole Body
Yoke

Section.

f?Jop_

S5j30d

LanKth

_4j?
4 '2

.Volts.

B. I A.T.P

A.T.

ZS^Qmo\\,%.

A.T.MMniAT.

t650a 52_
/4,0dd /4 5
5100

30
220

60
(070

/4-ld

EFFICIENCY.

Friction and W—
Iron Loss

\\\ load.

i^eld^ Loss RQtpx_

Arm. &c. I'R

Bnish Loss

Output
Input

Efficiency /a-

"26

?3^

73_
7400
I473

Fall.

_/7

H2X>
//76

954

17
75

'/O

,3.

42-5 \33 -6^ 29

jS40_ \S60 \260

883 \594 \309

Volts.

A.T.P

A.T.

Conr^mutator.

—Speed

Dia.„

Bars , .

Volts p. Bar-

Brs. p. Arm
Size of Brs.
Amps p. sq.
Brush Loss
Watts p. Sq.

Mag. Cur. P^ Loss Cur. 5
Perm. Stat. SIol /'P3

. , Rot. Slot x^= 1-96

.. Zjg-zag
z y.47 x5 73
1-77 X 5^0 X 3

/•64

Em\2-45 k43S X /2
/330 Amps : Tot
r='067 : X.

-2S60
-/275

^/35

- /'3f

2^2\^_3js90'5\ S, 1^.2-4^ ■ r^=..074 +0766

Imp. \-023-\-/'69 = /•3/
Sh. cir. Cur. .. , /320

Starting Torque O3€offutl
Max. Torque 2 -7^ times
Max. H.P 2 5 times

Slip 1:2^ %

Power Factor 0'3&

INDUCTION MOTORS 449

In the case of the 1500 h.p. motor, as a rather heavy overload is required for
3 hours, it will be well to provide room for a little e2ctra copper. We will therefore
take Kq at about 2. The diameter may be varied over fairly wide limits without
appreciably altering the cost of the motor. A speed of 6000 feet per minute is a
very suitable speed, and gives an economical motor where the frequency is 50 and
the motor of large size. 6000 feet per minute gives 2 feet per cycle ; that is to
^Ay> it gives a pole pitch of 12 inches.

J, 6000x12 ^_, ^„^ _ -^, mxEp,n

D^ jTTTT-x 2-54 = 239 cms., 5x 10^ = ^^)

IT X 246 ' K.v.A.

. , 5x10^x1360 ,^.

. . ( = ^TSTj TTTTT ^-77; = 4o'0 cms.

239 X 239 X 246

A preliminary calculation based on these figures will show us that 48'5 cms.
can be reduced to 47 cms. without unduly saturating the iron of the teeth.

In order to fix upon a suitable number of conductors, we want first to find
approximately the magnetic loading Afi of the frame. Now, it will be found that
in large induction motors the most suitable maximum flux-density in the gap
is about 6000 G.G.S. lines per sq. cm. A higher density in the gap up to 9000 c.G.S.
lines may be employed in " iron " motors ; that is to say, motors with wide teeth
and restricted copper space. Such motors call for a large magnetizing current,
especially where the air-gap must be fairly great. As the air-gap must be reasonably
great (say 2 mms.) on a motor of large diameter, the magnetizing current would
be too great if the flux-density were much greater than 6000, and the magnetic
pull would be too great if we were to reduce the air-gap.

Taking Bmax at 6000, provisionally, we get

^^B=750 X 47 X 6000=211 x 10«.

From formula (1), page 24,

3000 volts = 0-41 X «^ X Za X 2-11,
Za = 835.

The number of poles will be 24, giving a synchronous speed of 250, and a slip
of 1'5% will give a full-load speed of 246.

The number of conductors Za should be divisible by 24 and again by 3. The
nearest number to 835 which satisfies this condition is 864=24x12x3.

If, then, we have 12 slots per pole,* and 3 conductors per slot, the arrangement
will be suitable. This gives us 288 slots in the stator. In the choice of the number
of slots, regard must be had to the number of segments of stampings which make
up a complete circle. There should be, if possible, an even number of slots per
segment. As it is desirable, where possible, to have the standard number of seg-
ments divisible by 6, we might in this case have 6 segments.

Now, work out the actual AgB with 864 conductors

_3000o<iO«__
"^^^"0-41x4-16x864~'-"*''^^-

Take a calculation sheet (see page 448) and fill in the preliminary data.

The drawings of the motor are given in Figs. 406 to 409.

* For the considerations which settle the number of slots per phase per pole see pages 422
and 320»

w.M. 2 F

450

DYNAMO-ELECTRIC MACHINERY

lichca. e 9 6 3 0

hltiihiliil

5 FeeC.

Fig. 407. — Sectional drawings of a 1500 b.p. induction motor.

INDUOriON MOTORS

Flo. 408,— Bind of rtator aod rotor : i (nil alia.

dealgned lo meet Spadflcatlon No. 7, p. 138. Scale 1 : H,

452

DYNAMO-ELECTRIC MACHINERY

In order to reduce the unbalanced magnetic pull to a minimum, it is a good
plan,, on a large motor of this kind, to wind two paths in the stator in parallel.
This is comparatively easy to do on a 3000 volt motor. Instead of making 3 con-
ductors per slot, we can make 6, and divide each of the phases into two paths in
parallel in the manner indicated in Fig. 409. There, phase A is divided into two paths
A I and A^ which lie on opposite halves of the frame. It is impossible for the flux
on these two opposite halves to be very unequal, as that would necessitate unequal

Fio 400 —Diagram of winding of atator, abowlng two paths in parallel in each phase, to

minimise the unbalanoed magnetic pnJl.

back electromotive forces on the two windings A^ and A^^ The diameters of the
frame which divide the phases B and C are set at angles of 120^ to the diameter
which divides phase il, so as to ensure an equal distribution of flux, notwith-
standing a displacement of the rotor in any direction.

We thus get 131 amperes per conductor and 262 amperes per terminal.

The fixing upon the number of slots per phase per pole turns upon such con-
siderations as the following : The more slots we have the less will be the leakage,
and the better the cooling of the armature coils. As each armature coil must be
fully insulated to earth, a large number of slots per phase per pole would take up

INDUCTION MOTORS 463

a great deal of room, particularly in high-voltage motors. It thus comes about
that in very high-voltage motors the number of slots is kept down to the lowest
minimum, consistent with proper cooling and a sufficiently low leakage. It is not
advisable to have less than two slots per phase per pole, and if the total current
per slot is made very great (say over 1500 amperes), it becomes difficult to conduct
through the insulation the heat generated in the coil. In all cases where the proposed
arrangement involves a rather large total current per slot, a calculation should
be made of the difference of temperature between the inside and the outside of
the insulation in the manner indicated in the example given on page 224. Where
the voltage is low and the insulation space required not excessive, the number
of slots per phase per pole can be made greater and the stator leakage
decreased.

In the case under consideration the choice of 4 slots per phase per pole
gives a slot pitch of 2*6 cms., quite a reasonably large value for a 3000-volt
motor.

In provisionally deciding upon the size of conductor to employ, one may be
guided by considerations of current density in the copper, but the final choice of
size must depend upon the cooling conditions. A conductor 0*37 sq. cm. in section
will carry 131 amperes at a current density of 354 amperes per sq. cm., which,
having regard to the over-load conditions, appears to be a reasonable figure. The
six conductors, each 0*38 x 10 cm., are best arranged as shown in Fig. 408a, with
a spacer of micanite 0*5 mm. thick between each conductor, held in position with
half-lapped tape around each conductor. The allowance to make for this style
of insulation is 13 mms. per conductor. The whole coil is insulated with paper
and mica wrapping, and finally taped to a total thickness of 2*2 mms. After making
allowance for wedges and clearances in the slot, it will be found that a slot measuring
1'5 cms. wide by 4*2 cms. deep will be sufficiently large. It is always well to specify
plenty of room in the depth, as the cost of the machine is only very little increased
by so doing, and the little room in the depth greatly helps in the getting in of a
coil that is rather tight in the width. In width the coils should be designed to be
a reasonably good fit, so that the heat may be readily conducted from the insulation
to the iron, and so that the coil shall not vibrate in the slot.

Before the size of the slot is finally fixed, it is necessary to find the saturation
in the teeth. The cross-section of all the teeth is foimd exactly as described on
pages 71 and 322. The figures are given in the calculation sheet on page 448. The
axial length of the iron should be adjusted, so that the flux-density in the teeth,
one-quarter of a tooth length from the narrowest part, is not more than 17,000. In
this case, with a core length of 47 cms. and 7 vents, each 0*8 cm. wide, the flux-
density comes out 16,900. The maximum flux-density allowable depends partly on
the length of the teeth (for with long teeth the magnetizing ampere-tums become
excessive if the flux-density is high), and partly upon the permissible iron loss.
With teeth not deeper than 5 cms. and a frequency of 50 cycles, we may take 17,000
lines per sq. cm. as a suitable figure, where the allowable temperature rise is 40° C.
For a temperature rise of 45® C. we might allow 18,000. At 25 cycles one will go
up to 19,000, except in cases where it is necessary to keep down the magnetizing
current to the lowest possible value.

454 DYNAMO-ELECTRIC MACHINERY

The final arrangement, then, is 288 slots in the stator of the size shown on the
calculation sheet, 6 conductors per slot (1728 in all), with two paths in parallel,
giving virtually 3 conductors per slot. These figures are entered on the calculation
sheet in the manner shown.

The measurement of the mean length of a conductor is best carried out on the
drawing of a similar machine. From Fig. 407 we get the length in the slot 47 cms.
and the length outside the slot 65 cms., giving a total length of 112 cms. Multiply-
ing 1728 by 112 and dividing by 100, we get 1930 m. for the total length. Multi-
plying the constant 890 (the weight in kilograms of 1000 metres of conductor
1 sq. cm. in cross-section) by 0*37 sq. cm., we get 330 kilograms for the weight
per 1000 metres of conductor, and multiplying by 1930 we get 640 kilograms
as the total weight of the stator copper. This gives us 0474 kilogram per K.V.A.,
not an excessive figure considering the operating conditions.

The resistance of the winding we find as on page 143. 017 divided by 0*37
gives 0-46 ohm as the resistance of 1000 metres of conductor.

0*46 xl'93 =0*89 ohm for all conductors in series.

0*89/4 » 0*222 ohm, two paths in parallel ; 0*074 ohm per phase.

To find the difference in temperature between the inside and the outside of
the insulation, find the watts per sq. cm. of cooling surface. One metre length of
coil will have a loss in it of

131 X 131 X 000046 x 1*2 x 6 = 58 watts.

As the mean perimeter of a coU is 104 cms., the surface may be taken at 1040

sq. cms. The watts per sq. cm. are 0054. As the thickness of the insulation is

0-25 cm. and the conductivity 0*0012, we have, according to the method given

on page 222,

0*054 X 0-25 ,, ^ori ^•o' t *. 4.

— r^.r,r^^n — = Hw^'C. difference of temperature.

Thus, if the iron of the teeth is 35® C. above the surroimding air, the copper
in the slots will be about 47° C. above the air. The allowance of 1/0054 = 18*5 sq.
cms. per watt for the exterior of the stator coils, which are subjected to a good
draught from the rotor, will ensure the temperature rise of the ends of the coils
being well below 40° C. rise (see page 324).

The methods of calculating the cooling surfaces of the stator and the rate at
which heat is given off from them are the same as given on page 325 in connection
with the 750 k.v.a. generator. The figures are given on the calculation sheet
(page 448). The total losses to be carried away from the stator surfaces are, on a
liberal computation, 24,300 watts, and with 40° C. rise of the frame we can get
rid of 25,700 watts.

The rotor winding. In fixing upon the size and number of rotor conductors,
the first step is to decide upon the voltage to be generated in the winding when
the slip is equal to the synchronous speed. In large motors we will make this as
high as is consistent with safe operation. The higher the voltage the less the
current, and the less elaborate the brush gear for collecting it. If the voltage is
made too high, it may be dangerous to persons starting the motor, if the brush
gear is not perfectly protected ; and, moreover, the insulation of the winding is

INDUCTION MOTORS 455

more difficult to carry out for a high voltage. For large motors of 1000 h.p. or
more, rotor voltages of 800 to 1000 are common, and there seems to be no objection
to rather higher voltages for very large motors where it is worth while to com-
pletely protect the brush gear. In the motor under consideration, if we make
15 slots per pole on the rotor, and use a barrel winding with two bars per slot,
the ratio of transformation between stator and rotor will be

288 X 3 ^ 864
360 X 2 720'

If all the rotor conductors of one phase were put in series and connected in
atar, the voltage on the collecting rings at the instant of starting up would be

720

3000x^=2500.

864

If the phases were connected in delta, the voltage would be

2500

1-73

= 1445.

If the conductors were connected with two paths in parallel and in star, the

2500
voltage would be — jr- = 1250. The latter seems a suitable voltage for so large a

motor, and if this arrangement be adopted, the current per ring at full load will be

1500 X 746 KOA
1250x173°^^ amperes per rmg.

The size of the conductor will then depend upon the amount of slip which we
wish to have at full load. If we want to have only 1^ per cent, slip at full load,
the resistance of the rotor winding must be adjusted so that the I^R losses in the
rotor are \\ per cent, of the input to the rotor, or approximately IJ per cent, of
1120 K.w. ; that is to say, 14 K.w. Having fixed the voltage of the rotor winding,
and therefore the rotor current at fall load, the resistance per phase to give any
percentage loss is easily calculated. Allow 1 K.w. for losses on contacts, etc., leaving
us 13 K.w. on the winding itself, or 4340 watts per phase.

4340= 520 X 520 xr„

f2=0016 ohm hot, or, say 0-0133 ohm cold, per phase,

with two paths in parallel, or 0-16 ohm with all conductors in series.

Now the length of one conductor is 95 cms. So the length of 720 conductors is

685 m. If this length is to have a resistance of 016 ohm, the resistance for

1000 metres will be 0-234. If we choose a conductor measuring 0-5 cms. x 1-5 cm.

0-17
and having an area of 0-73 sq. cm., this will have a resistance of ^^wo =0-234 ohm

per 1000 metres. This will give a resistance per phase with two paths in parallel
of 00133 ohm (cold), or 0016 ohm (hot).

It is convenient for many purposes to transform the resistance of the rotor

winding by multiplying it by the square of the ratio -^ . We are then able to add it
to the stator resistance in calculations of effects occurring in the stator which

456 DYNAMO-ELECTRIC MACHINERY

depend upon the rotor and stator resistances. This transformed resistance of the

rotor we will denote by fj.i. In the present case the ^* = ocH' *^°d the square of

this ratio is 5*76. '

rj. 1 =0-016 X 5-76 =0092 ohm (hot),

while ri=0089ohm(hot).

Therefore ta, the apparent resistance of the motor per phase to alternating
current applied to the terminals of the stator is ri + r2.i =0-181 ohm (hot), or
01506 ohm (cold).

The rotor winding consists of two conductors per slot. The insulation around
each conductor is paper and mica and tape, to a total thickness of 1 -8 mm. The
whole, with a suitable slot lining and wedge, will go in a slot 0*96 cm. wide by
4-2 cms. deep, as shown in Fig. 408.

The calculation of the flux-density on the teeth is carried out as shown on page
448. The figures are given on the calculation sheet.

The next step is to calculate the magnetizing current. This depends mainly
upon the length of the air-gap. It is not advisable to reduce the air-gap of a large
motor of this kind much below 2 mm. (see Fig. 401). This air-gap is perfectly
satisfactory if provision is made for neutralizing the unbalanced magnetic pidl in
the manner explained above, and if the design and workmanship on the stator and
rotor frames is good.

The contraction ratio when worked out in the manner indicated on page 417 is
found to be 118. The maximum flux-density in the gap is

204x108 ^j^,.

A.T. on the gap =0-2 x 1 18 x 5800 x ^^ = 1090 ampere-turns.

The magnetizing current. The calculation of the ampere-turns on the stator
and rotor cores and teeth is carried out as indicated on page 448. The figures are
given on the calculation sheet. The total ampere-turns per pole are 1410.

To get the magnetizing amperes per phase Im we adopt the rule given on page
420. For a three-phase, star-connected, full-pitch winding we have

0 -437 X I„,Za 0 -437 x /,„ x 864

= 1410.

poles 24

Im = 90 amps.

That part of the no-load current which is in phase with the voltage is obtained
by dividing the no-load watts by the voltage and by 1-73. The iron loss in this
case amounts to 16-8 K.w., and the friction may be taken at 8 K.w., giving a total
no-load loss of 24 -8 k.w. Thus the watt component of the no-load current amounts
to 5 amps. If we take 0 for the centre of our clock diagram, as in Fig. 410, we
can set off ON' to scale to represent 90 amps., and N'N to represent 5 amps.

The next step is to calculate the short-circuit current. As stated above, the
most accurate way of arriving at this is to rely upon tests of similar motors built
on the same frames or on similar frames. If, however, no such data are available^
we may calculate the value of the short-circuit current with a fair approximation
by the use of the rules given above for the calculation of the slot leakage, the zigzag

INDUCTION MOTORS 457

leakage and the end leakage. It should be pointed out, however, that these rules
do not take into account the saturation of the iron along the leakage paths, which
will probably occur before the current reaches its full short-circuit value. This
saturation, however, is only of importance when we wish to know what the actual
starting current is. The power factor of the motor, and other particulars of its
performance at normal load up to two or three times full load, will be dependent
upon the diameter of the circle constructed by taking for the value of the short-
circuit current the value that it would be if there were no saturation. The actual
starting torque of the motor, however, is dependent upon the amount of saturation
which occurs on short circuit. It cannot be determined with any accuracy by
calculation. Indeed, two motors built from the same drawings give different
starting torques, depending upon slight difierences in the amounts of iron in the
armatures.

The methods of working out the leakage in stator and rotor and the end leakage
have been given on pages 420 to 427. We found that the total leakage per pole
for one ampere per phase in the stator winding amounts to 4135 c.G.s. lines. The
working iSuz per pole is found from the formula :

A,BxKf _2»04xl08xO-66^g^^.^y
No. of poles 24

Therefore the short-circuit current, if there were no resistance, would be

c^/> 5-6xl0«

And we have seen on page 428 that the apparent impedance of the stator is
1 *31 per phase, so that actual current on short circuit is

Eg 1730

r""i-3i

We must next calculate the watt component of the short-circuit current. The
resistance of the stator per phase is 0'074. The actual resistance of the rotor per
phase is 0*016 ; but, the ratio of transformation being 864 divided by 360, or 2*4,
we must multiply 0*016 by (24)2 ^ reduce the rotor resistance to its equivalent
for a one-to-one ratio. This gives us r2,i=00766. Therefore fi + r^i^Olb (cold)
or 0-18 (hot) per phase. Multiplying by the square of the short-circuit current,
and by 3, we get :

0 18 X 1320 X 1320 x 3 = 940 watts loss on short-circuit.

Dividing this by 3000 volts and 1 '73, we arrive at 180 amperes per phase for the
watt component of the short-circuit current. Referring now to Fig. 410, we set
off the 180 amperes shown at FS and 1320, represented by O'S. The usual practice
is to place the point 0' at a position midway between the points 0 and N ; the
reason for this is that the magnetizing current on short circuit is reduced to about
half its normal value, so that the point 0 really moves half-way towards N. We
now know that N and S lie upon the semicircle of the Heyland diagram. The
centre of the semicircle is now found by the construction given in Fig. 400 (page
411) and the semicircle drawn through N and S, It will be observed that OS is
the short-circuit current calculated on the assumption that there is no saturation

j„„^=.i!f^ = 1320.

458

DYNAMO-ELECTRIC MACHINERY

in the path of the leakage line. The semicircle which we have drawn is the locus of
the point P of the radius vector OP, which represents the stator current for all normal
loads. For greater stator current we should expect some saturation to occur in the
leakage path, and the point P will then follow a locus such as that given by the
dotted line in Fig. 415, the starting current being somewhat greater than that

O 0* A/'

Fio. 410. — Circle diagram of 1500 H.P. motor. Scale 1 mm.»10 amperes per phase.

given by OS. The load line is given by NS. The usual method of obtaining the
torque line NT \& to divide 8F in such a manner that

ST_^r^
TF" r^'

As, however, the short-circuit current in the stator is somewhat greater than that
in the rotor, it is better to divide SF so that

8T jr^^xNS^
TF~ r^xO'S^'

This is done in Fig. 410. We then know that any perpendicular such as PX drawn
from the semicircle to the load line represents the watt component of the stator
current, which when multiplied by the volts and 1-73 gives us the output of the
rotor. The value of PX for full load can therefore be found by dividing the full-load
output, 1120 K.W., by 3000 and 1-73, giving us 216 amperes per phase. Setting
up a perpendicular 21*6 mm., we obtain the full-load stator current OP and the
full-load rotor current NP, The power factor, which is the cosine of the angle <^,
is found to be 0*89. The maximum torque is found by drawing a tangent through
Q parallel to the torque line NT. The vertical line QR intercepted by the torque
line, when scaled off, gives us 570 amperes per phase, and this at full speed would
be equivalent to an output of 2960 K.w. The maximum output is obtained by
drawing a tangent through U parallel to NS. The perpendicular UW, when scaled
off, gives us a stator current of 530 amperes per phase, which would give a mATiTmim

INDUCTION MOTORS 459

output of 2700 K.W. The efficiency of the motor is worked out from the separate
losses, as indicated on the calculation sheet on page 448.

To obtain the slip we must find the ratio of the rotor losses to the rotor input.
The actual current in the rotor is obtained by scaling off NP in Fig. 410 and

multiplying by ^, We thus get

217 X 2 -4=: 520 amperes per phase.

Each of the three phases has a resistance of 0-0133 (cold) or 0016 (hot), so that
the I^R losses at full load are

520 X 520 X 0-016 x 3 = 13,000 watts.

To this loss we should add about 1 K.w. for brush losses. The input to the rotor

14
will be 1120 + 14 = 1134 K.W., so that the slip = =^^=0-0125, or 1-25 per cent.

460

DYNAMO-ELECTRIC MACHINERY

SPECIFICATION No. 8.

ChAracterfstics
ofiMotor.

350 H.P. INDUCTION MOTOR FOR PUMP DRIVING.

120. The Contractor shall supply and erect as described
below an induction motor having the following characteristics :

Normal output

Normal voltage at ter-
minals

Frequency

Number of phases

Speed

Power factor not less
than *

How connected to load Direct-connected to centrifugal

pump.

Temperature rise after
6 hours full-load run 45° C. by thermometer.

350 H.P.

2200 volts.

50.

1350 to 1480 R.P.M.

Over load
Temperature rise after

3 hours 10 per cent.

over load
Maximum torque
Starting

Puncture test

10 per cent, for 3 hours.

55° C. by thermometer.

1-5 times full-load torque.

By means of rheostat in rotor

circuit.
5000 volts on stator.
2500 volts on rotor.

Nature of
Load.

121. The motor is for the purpose of driving a centrifugal
pump, and for this purpose shall be direct connected to tiie
shaft of the pump situated near the sump well of a coal

mme.

Variation of
Speed.

122. The normal speed of the pump is 1475 r.p.m., but on
certain occasions the speed must be reduced, and may then be
between 1350 and 1475 r.p.m. For giving this range of
speed, and also for starting the motor, the Contractor shall
supply a metallic rheostat fitted with a suitable dial plate
having not less than 10 steps.

* The Contractor is to state the power factor at full load and three-quarter load of
the motor he proposes to supply.

THE SPECIFICATION OF INDUCTION MOTORS 461

123. The Contractor shall deliver a separate quotation for s^^^^^
this rheostat, giving fall particulars of its construction and

the temperatuje rise guaranteed after a 6 hours' run at 300 h.p.
at 1350 R.P.M.

124. The motor will be situated in a dry chamber and situation,
supplied with cool dry air. It must be suited in every way

for the class of work for which it is intended.

125. Some of the gangways leading to the point where the £J^*y «'
motor is to be erected are not more than four feet high by

six feet wide. The dimensions of the motor and its supports
must be such that it can be taken along the said gangways.

126. Plan of the mine and of the proposed place of erection Pun.
can be inspected at the offices of the Purchaser.

127. The bedplate will be supplied by the pump makers. Bedpute.
The motor shall be supplied with such feet or other supports

as shaD be suitable for fitting and bolting to it. Particulars

of this bedplate and the height of the running centres wiU be

suppKed to the Contractor within three weeks from the

civing of the order. At the same time, the Contractor will niu coupling.

be given the dimensions of the half coupling, which is also

to be supplied by the makers of the pump.

128. The Purchaser will undertake the lowering of tl^^g^J^^jj^
motor into the mine and the conveyance along the gangways, Mine.
provided he is satisfied that the outlines of the motor make

it possible ; but the Contractor shaD carry out the erection
and setting to work of the motor.

(See Claufles 99, p. 441 ; 135 to 137, p. 460.) Efficiency.

(See Clause 101, p. 441.) £l2S,f*wori„.

129. The tests taken at the maker's works having been Testa on
carried out satisfactorily, the efl&ciency of the motor shall '

be taken as proved.*

* In caaes where one Contractor makes himself responsible for the whole plant,
pump and motor, it is usual to ask for a guarantee of the combined efficiency of the
plant. This is sometimes expressed in terms of so many gallons of water per hour
raised a certain height for the consumption of so many electrical units. In asking for
guarantees of this sort, care should be taken to specify exactly the points between
which the head of water is to be measured. Where tests are to be earned out, it must
be clearly stated which party shall provide the measuring tanks and bear the cost of
the t^ts.

462 DYNAMO-ELECTRIC MACHINERY

Tastof Power . 130. After the plant is installed, the Purchaser may call
^ '■ for a test of the power factor of the motor when running

under the conditions specified in the guarantee. Such test
shall be carried out by taking the ratio of the two readings
of a cahbrated wattmeter connected first in one phase and
then in another. The power factor as worked out from these
two readings shall be taken to be the power factor of the
motor.

instniments. 131. The Coutractor shall supply all instruments for this

purpose. The cost of recalibrating instruments shall be
borne by the party requiring the same, unless the instrument
shall be proved to be 1 per cent, out of calibration, in which
case the cost shaD be borne by the Contractor.

Puncture (See Clauses 318, p. 611 ; 234, p. 564.)

DESIGN OF A 350 H.P., a-PHASE INDUCTION MOTOR TO COMPLY

WITH SPECIFICATION NO. 8.

2200 volts ; 50 cycles ; speed 1350-1475 R.P.M.

As this motor is for the purpose of driving a centrifugal pump, it is not necessary
to give it a very great over-load capacity. A maximum torque equal to 1 -5 or 1 -7
of the fall-load torque will be quite sufficient for the purpose, tinder the circum-
stances, an output coefficient, Kq=2'6 (see page 447) will be ample. This gives
us a D^l constant of 4 x 10^. The ratio between diameter and length might be
varied over fairly wide limits without appreciably afiecting the cost ; and it wiU
be impossible, without going very closely into the cost of labour and material
in any particular factory, to decide which ratio is best. The speed of the motor
being high, the designer will avoid making the radius too great. A diameter of
46 cms. will give a peripheral speed of 36 metres per second, a speed sufficiently
high for good ventilation and yet not excessive. With D=46, we get 1=38-5,
and this ratio will be found to be very economical. It will be noted that the
specification requires a motor, the outside diameter of whose frame is less than
four feet.

The principal steps in the calculation of the motor will be seen from the
calculation form on page 463. Drawings of the motor are given in Figs. 411
and 412.

This is essentially a " copper " machine. Having 4 poles of wide pole pitch,
we would, with normal proportions of copper and iron, obtain a motor with a
maximum torque some three times full-load torque, a ratio much greater than is
needed in this case. Moreover, if the cross-section of copper were kept down the
normal rating of the motor would not be so high as it is. In the design given,
the copper section has been increased at the expense of the iron section, until the
motor has a maximum load not more than 1 -7 times its normal load* In making the

INDUCTION MOTORS

463

U^itQJt'Aug'i^/S.. Type.... .•eeit...../^8YN MOTOR

KVA.30S..\ P.¥:SZ\ Ph*sc3 ; Vo\t»22Q0. ; Amp«perter..^^.

RP.A^^ ..Amps p. cond. B.Q. Amps p. br. arm. Temp, rise ^$ ,.Q.

.^.... .Poles Elec. Spec ...^...

CjtXtt.SQ. ...I R.P.uMJQ.Q..', Rotor Ampe
..Reflation Oreiioad

Customer : Order

Frame 7^.

Air

No ; Qttot No : Perf. Spec .^.', Fly-wheel effect

t'5 Circum /44. . . ; Gap hsta.^'^^Q.

posa. Ag B

A,B.?-?f<?.^/^.*

; poss.

67^(>0. ;cS5;ii.T^P.

D* &.XRPM

K.V.A.

4-i(f0^

K^-4/5 ; M^Q Voly ='^/5« .2y.. X !9.4<> x TJ^^i^

Arm. A.T. p. pole Mais. Fid. A.T.

Armature, ftev.

Stat.

Dia. Outs

Dia. Ins . —

Gross Length

Air Vents

Opening Min _Mean

•fM Velocity

Net Length.5:i:5 X 89

Depth b. Slots ._

Section - 3f6 , Vol

Flux Density

Loss-^2ipcu.QZL„ Total
Buried Cu./fl5i? Total
Gap Area i^^i? . Wts
VentAreay6^^C?.:\Vts
Outs, hx^ f 4,000 .y\\s

75 _
'_ 46'

4 — 7S

' iMu Circ
6p\>,f'7 =

No of Segs
No of Slots

Section Teeth

Volume Teeth

Flux Density . ^

Loss-/? p. cu C/7? Total

36mper§ec._
31 6 \

10 i_:i_

e^Aoo^ - -

6700\

/600 __

2^50 4:300 _
2000,^
1200 \
2100 \^3QP

1^ -

/49
102

I t4-90
I 7060
/6/MO

esa

Weight ol Iron.

560

Star nr Mwti -.Throw

Cond. p Slot -

Total Conds - .

Size of Cond. '^^ x >54
.\mp. p. sq C^. '

Length in Slots 39 I
Length outside ^<^_Sum
Total Length S9 _
Wt. of 1. 000 205 Total
Res p 1. 000 '?^. Total

Watts p. /Z7-

Surface p. m. — _

Watts p. Sq.

•08 kO^

— i^um

i'20,2-)9,3f6&c

.- '-^-1

Q^O I . ._.
• 23 sq cm.

'3^0 ;_ V

^640 m

/70J(gs.
62 J -24
___ &0.
WOO

'08_
13 3 "*

M Slots

^/'7i>
>

■«-

\€---24- ■-■*

y.>e

/S --»

I •

< 9

; ^ S^6 Slots

'/2SX/23 X 44S0X'jg6 « 545

'rield Otafc oi^ Rotor.

Dia.

\ Total Air Gap

Gap Co-cff. K.

Pole PitcluSS. Pole
K,

Arc

Flux per Polc^lSULZfif.

Leakage n.l f.l

Area Flux density -

Unbalanced Pull L_

No. of Seg
No.ofSla
Vents-
K.

.Section-Zf_^<?-.

43-75

125

1-23

132

n.4-oo

Weight of Iron,

Shunt.

27Q^

SOflM.

Kacorwaa.

A.T.pPolen.Load-
AT p. Pole f. Load _
Surface pvimattlt.
Surface p. Watt.-

I- R.

I R.

Amps.

No. of Turns

Mean 1. Turn __

Total Length

Resistance

Res. pen. 000

Size of Cund

jm2_\

Conds. per Slot

Total

Length

Wt. per 1,000

Total Wt.

Watts per Sq

Star or Mesh

Paths in par.ille]

O^i

0-2^

7X/0
2__
792

'016 p^rph

Comm.

lSZ-

92X192 ^/76m.

630

107 kffs

Star

Magnetization Curve.

Core

Stator Teeth
Rotor Teeth

G.\p

Pole Body -
Yoke

Section Length

(49(1 44__
I4l0j Zl5
5550 J25

Volts.

iilT.P

A.T.

^/.-f<?. Volts.

A.T.P<^JAT.

f^^oa ^5J 2oo_

f7.4-oo 7^' ^2^
^MP_ ___ 545

! 20

960

trnciENCY.

Friction and W
Iron Loss _ _
♦^ieH Loss in Rotor
Arm. &c. PR
Brush Loss

,4 load. Full.
2-5

_±
1-9

2:AJ

Output

luput ^

ElViciency f/Q-

Volts.

A.T.P i A.T

Commutator.

Dia Speed —

Bars

Volts p. Bar^

Brs. .p. Arm

Size of Brs.

Amps p. sq.

Brush Loss ,

Watts p. Sq

Mag. Cur. /^ 4 Loss Cur. //^
Penu. Stat. Slot f5J

.. Rot. Slot x^ = -9
.. Zig-zag ^^ 2 2

2 X38 5x4.€/'3SS

' 77X 356 >:/4 =6S20
E'.\il2-4jA 4-6 X 70 ---- 7900
256 .\mpf ; Tot /6, 720
^ = '0407 . X. = J

^^1^^4 36 .T^^ '24 +-5/

+ 25

^2&L

S03

Imp. ^/^3

Sh. cir. Cur.

Starting Torque '22gffull

Max. Torque /•? times

Max. H.P / '62 times

Slip ''^5%,

Power Factor 0'9!L

464 DYNAMO-ELECTRIC MACHINERY

calculation, tke method adopted is essentially the same as that given on pages
445 to 459. We have only four poles, and are able to have as many as fifteen
slots per pole. We choose a very high current loading per cm. of periphery. Even
with a figure as high as 466 amperes per cm., the temperature will not be too high
where very efficient ventilation is adopted. We may therefore choose as many as 840
conductors, and in so doing we cut down the value of AgB^ and with it the magnetiz-
ing current. Allowing 2\ per cent, for the drop in the stator winding, the back
voltage generated by the revolving field may be taken at 2140 volts. We then have
the formula 2140 =0415 x 25 x 840 x AnB,

^^8 = 2-46x108.

We have in the stator 60 slots, and 14 conductors per slot. Working the copper
at 350 amperes per sq. cm., we adopt a conductor 0-42 x 0-5 cm. Fourteen conduc-
tors with insulation (see page 202) will occupy a slot 1-7 x 4-3 cms. The diameter
of the mean circle through the slots will be 149 ; subtracting from this 60 x 1 -7 = 102,
we get 47 cms. for the total width of the teeth. The net length of iron is 31 -6 ;
the cross-section of all the teeth 1490 sq. cms. Dividing this into 0-246x10^,
we get a flux-density in the teeth of 16,400. From Fig. 29 we find that the loss is
0 12 watt per cu.* cm., giving a total loss in the teeth of 850 watts.

The flux per pole is obtained from the formula (1) on p. 326.

0-246 X 108x0-7 , o 1^ 1- 1
— - — = 4-3 X 10" C.G.s. lines per pole.

Allowing 10 cms. depth behind the slots, we get a cross-section of core of
316 sq. cms.: ^,^^^^

-s — ;rT7r- = 6700 C.G.S. Uucs per sq. cm. in the core.
2x316 ^ ^

This gives a loss of 0 025 watt per cu. cm., and a total iron loss of 2450 watts.
The buried copper loss amounts to 1850, giving a total of 4300 watts, to be dissipated
by the iron surfaces of the core. It will be seen that with 45° C. temperature rise,
the iron surfaces of the core can dissipate 5300 watts. We find that there are 80
watts lost per metre length of armature coil. As the cooling surface per metre is
1000 sq. cms., we have 0 08 watt per sq. cm. As the thickness of insulation is
0 02, and the heat conductivity 0 0012 (see page 221), we may eicpect a difference
of temperature of 13-3° C. between the copper and the iron.

Flux-density in the air-gap. This is obtained by dividing the AgQ by the

«*Par«a- . 0-246x108 ,,^^ „

5550 =^^50 = B,.

The gap coefficient, worked out from Figs. 36 and 37, is 1 -23 ; so that the
ampere-turns on the gap are :

0125 X 1 -23 X 4450 X 0-796=545 ampere-turns.

The flux-density in the rotor teeth, as worked out on the right-hand side of
the form, is 17,400.

Magnetizing current. We now proceed to calculate the magnetizing current.
We see from the sheet that at 2140 volts the ampere-turns per pole required are 960.

INDUCTION MOTORS

I" I

II i

at I

i 3

t 5

466 DYNAMO-ELECTRIC MACHINERY

As there are 210 conductors per pole, the magnetizing current is obtained from the
formula: 960

'^'» = b437x210 = ^^-^''°'P*'^-

The core-loss current equals 2450. To find the current in phase with the voltage
at no load, we add the iron loss 2450 watts to the friction and windage losses, which
may be estimated at 1900, and divide the sum by 2200 x 1 -73. This gives us
1-15 amperes.

Botor conductors. The output of the rotor is 350 h.p., or 261 k.w. A standstill
voltage of 500 will give us about 300 amperes per ring. This is a suitable current
for a motor of this size. To generate a standstill voltage of 500, we must have a
transformation ratio of about 4-4. If we choose 96 slots and 2 conductors per
slot, making 192 conductors in all, the transformation ratio will be

192 *'^'

We therefore decide upon 192 conductors. They will form a barrel winding,
as shown in Fig. 411.

A current density of 460 amperes per sq. cm. is not too high for the rotor copper,
so that a conductor 0-7 cm. will be large enough. Two conductors 0-7x1, with their
necessary insulation (see page 202), will occupy a slot 0-9 x 2-4 cms. The total
length of all the rotor conductors works out at 176 metres, and they have a total
resistance of 0 041 ohm ; so that the resistance per phase is 0 0137 ohm cold,
or 0-016 hot. Multipljring this by the square of the transformation ratio

^|^y=4-38», gives us r2.i = 0-31.

The method of working out the permeance of the stator and rotor slots and the
zigzag leakage is the same as that described on pages 422 to 427. The permeance
of the stator slots is 1 -51 ; for the rotor slots it is 1 -44, which, multiplied by the

ratio — , gives 0-9. The zigzag permeance works out at 2*2 ; so that the total
96

permeance for 1 cm. axial length is 4-61. This, multiplied by 38-5 and by 2, gives
us 356 for the permeance per pole. For one ampere passing in the stator winding
we have :

1 -257 X 1 -41 X 14 wires per slot x 356 = 8820 lines per pole for 1 ampere.

Next, consider the leakage around the end winding. Referring to Table XVIII.
page 427, the value for Kl for concentric stator winding and barrel rotor winding
is 2-45. The pitch of the pole is 36 cms., and the value of a„ is 10 cms., giving

Z + ao = 46.

There are 5 slots per phase per pole, each carrying 14 wires ; so that for one ampere
passing in the stator the virtual ampere-turns are 70. Thus we get the leakage
around the end wmdings equal to 7900, giving a total of 16,720 leakage lines per
pole for one ampere in the stator. Now, the working flux to generate full voltage
is 4-3 X 10* lines, so that it will take 256 amperes in the stator to produce enough
leakage to generate a back e.m.f. equal to the E.M.F. supplied. As the voltage
per phase is 1270, this divided by 256 gives us an apparent reactance of 5 ohms

THE SPECIFICATION OF INDUCTION MOTORS 467

per phase. In order to find more exactly the short-circuit current, it is necessary
to take into account the value of the stator and rotor resistances ; these are worked
out on the calculation sheet, r^, the sum of the stator resistance, and the rotor
resistance referred to the stator, is 0 -24 + 0 -31 = 0 -55. The impedance then works out
at 5*03 ohms, giving a short-circuit current of 253 amperes. The actual test on the
frame of this motor gave readings between 230 and 270 short-circuit amperes per
phase, depending on the position of the rotor slots relative to the stator slots.
From these data we draw the circle diagram as described on page 414. From it,
we find that the starting torque with no resistance inserted is 0-22 of the full-
load torque ; the maximum torque is 1 -7 times full-load torque, and the maximum
horse-power 1 -62 times the normal horse-power.

Slip. The slip is found by taking the ratio of the PR losses in the rotor at full
load, equal 4-4 K.w., to the total rotor input, equal 266 K.w.

4-4

^^ X 100 = 1 -65 per cent.

SMALL MOTORS.

In drawing a specification for a small motor, one should aim at making it as
simple as possible, confining oneself to those matters which are important from the
purchaser's point of view, and leaving to the manufacturer as free a hand as possible
in the design, so that he may be able to put forward one of his standard machines.
A standard motor will probably be much cheaper and more quickly delivered
than a special motor built to comply with a specification which too rigidly pre-
scribes its characteristics. It is particularly important that the specification should
be confined to performance, and not dictate the methods of manufacture by which
that performance can be obtained. It may be well in some cases to call for. a motor
having a certain power factor, but it is better to leave the efficiency to the manu-
facturer and see what figures can be guaranteed. The following form may be
taken as a guide in the case of a small motor.

468

DYNAMO-ELECTRIC MACHINERY

SPECIFICATION NO. 9.

35 H.P. INDUCTION MOTOR.

PnrpoBesof
the Motor.

132. There shall be supplied a three-phase induction motor
for the purpose of driving a Une of shafting in a carpenter's
shop.

Type of Rotor. 133. The rotor shall be of the squirrel-cage type,
charaoteristics. 134. The motor shall have the following characteristics :

Pulley and
Slide Bails.

Extent of
Work.

Normal output

Normal voltage at ter-
minals

Frequency

Number of phases

Speed

Power factor

How connected to load

Size of steel pulley to
be supplied

Temperature rise after
2 hours' fuU-load run

Over load

Maximum torque

Puncture test

35 H.p.

500 volts.

50 cycles.

960 revs, per minute.

Not less than 0*8.

Belted.

24" dia. x 12" face.

45° C. by thermometer.

20 per cent, for 15 minutes.

2'5 times full-load torque.

1500 volts alternating applied
for 1 minute between wind-
ings and frame.

135. The motor shall be provided with a pulley of the size
above specified, and be mounted on slide rails with belt-
tightening screws.

136. The contract includes the delivery of the motor at
the purchaser's works in , together with certain
switch gear and starting gear, but does not include erection
or starting-up.

THE SPECIFICATION OF INDUCTION MOTORS 469

137. The contractor shall state the amounts of the follow- Efficiency.
ing losses in the motor which he supplies :

1. Bearing friction and windage losses. (At no load.)

2. Iron losses at no load, when run on 500 volts 50 cycles.

3. Annature and rotor copper losses at full load, allowing
for temperature rise.

The contractor shall state what calculated efficiency he
guarantees on the basis of these separate losses, as weU as
the actual efficiency of the motor at full, three-quarter and
half load.

138. The motor shall be nm for one hour at full load at the TestB
contractor's works in the presence of the purchaser's engineer.
On this test the power factor shall be measured both by
power-factor meter and by the two-wattmeter method.
Measurements shall also be taken of the power taken to run
the motor at no load, and of the power absorbed when the
motor is locked and taking full-load current on short circuit.
Measurements shall be made of the maximum torque. When
the motor is still warm after these tests, a pressure of 1500
volts alternating, 50 cycles, shall be applied between the
stator copper and frame for one minute.

138a. If the motor is foimd to fulfil the guarantees, so far Acceptance,
as can be ascertained by these tests, it shall be accepted
without further tests. In view of the difficulty of measuring Efficiency,
the actual efficiency in a commercial test, the calculated
efficiency shall be taken as the criterion, unless there is very
positive evidence that the motor falls below its guarantees in
actual efficiency.

1386. If during the first six months after delivery any period of
defects in construction or performance become manifest, the m»'°*®'^®^-
same shall be immediately rectified by the contractor at his
expense. Any time elapsing between the reporting of the
defects and the remedying of the same shall not be coimted
in the six months' period of maintenance.

470 DYNAMO-ELECTRIC MACHINERY

DESIGN OF A 36 H.P. INDUCTION MOTOR.
500 volts ; S-phase ; 50 cycles ; 980 r.p.m.

A motor of this kind would form one of a manufacturer's standard line of motors.
Its rating would have been determined by actual trial, so that in practice one would
not work out its diameter and length from first principles, but take a motor from
the list whose rating is known. Nevertheless, it is of interest to apply the rules
which we have given above to see how far they are of use in predetermining the
performance of the motor from the dimensions.

In the first place, it will be found that for these small motors the output
coefficient, Kq, is smaller than in large motors. The output coefficient of an induc-
tion motor will depend upon the point of the circle diagram for the frame which is
taken as the full-load point. If a motor with great over-load capacity is wanted,
the full-load point (P, in Fig. 400) must be taken nearer the origin than where a
motor of smaller over- oad capacity is wanted, and the rating of the frame must
be correspondingly decreased. In this case the specification calls for a motor which
will yield 2*5 times full-load torque. Referring to Table XIX. page 447, we may
take Kq at something below 1*55. The D^l constant comes out at 6*5 x 10^. Take
a diameter 40 cms., and an axial length 14 cms. If a fairly good power ^tor is
desired on these small motors, it is well to keep the diameter great as compared
with the length, because on a great circumference one has more room for increasing
the number of slots, and the number of turns per pole can be increased, and thus
the magnetizing current will be decreased. The pole pitch is 20*8 cms., about 50 %
greater than the axial length, and the ratio of active length of stator conductor

14
to total length is only — . If the power factor were less important one could increase

the axial length of iron and reduce the diameter and reduce the weight of copper in
the stator.

The calculation sheet on page 471 gives full details of the steps in the working
out of the losses and cooling conditions. The motor is illustrated in Figs. 413 and
414. The circle diagram is plotted to scale in Fig. 415.

The large diameter gives us room for 72 slots of the dimensions given on the
calculation sheet. This is a convenient number for a standard motor, as it enables
the frame to be wound for either 6 poles or 8 poles.

The stator winding will be a '' mush " winding of the type illustrated in Fig. 138.
The number of conductors is determined by the magnetic loading of the frame.
With a standard punching and a given axial length, there is a maximum magnetic
loading beyond which we cannot go without saturating the iron too highly. We
may take a flux-density of 18,000 in the teeth as a suitable figure for these small
motors of 50 cycles. At 25 cycles we might go to a fiux-density in the teeth of
19,500, not merely because the iron loss is lower at 25 cycles, but because we have
fewer poles and consequently a larger number of ampere-turns per pole, so that we
can afford to have a higher magnetic reluctance. With 18,000 lines per sq. cm*
and 600 sq. cms. in all the teeth, we have an AgB of 0106 x 10®. Before we can
settle on the number of conductors, we must decide whether we will connect the

INDUCTION MOTORS

471

DaU ^^.s/PHfii^M^ Type ..«, OBIIb... ^SYN MOTOR ROVARY .^ .Poks EIcc Spec .-^

KVA. ^...,..; PF:>*?; PhueJ ; Voh»..^<?<? ; Amps per ter..3^. ; Cyd/tM.&Q. ...\ R-P-M.-S^.C?...; Rotor Amps-

H.P..^- Amps p. coad. ^^'^ Amps p. br. arm Temp. riM .4'S?..C. .Repiktieii Oirerload .ZQ.f9rJihr..

Costomcr i ; Order No ; Qnot No ; Pirf. Spec ;' PlT-wheel effect

Frame 5^^.

Air

CiTcom.m ;<^PAr«i^7'^^ :^B^/5«x/<?*:;:; I.

uz.

K, r^f. ; 4-9.Q ytoitt=--2* x t6:.6* US2 * JOa

Arm. A.T. p. pole....JW ft.O

IC.V.A.

66x/o^

Max. Fid. A.T.

Armature. «*v.

Stat.

o
O

I-

2
o

o
o

Dia. Outs.
Dia. Ins

Gross length

Air Vents none

Opening Min Mean

Air Vdocity

Net Length /4- x-89

Depth b. Slots

Section 7/'^ Vd.

Flux Density

Lom'gf p.cu.C^- Total
Buried Cu.37g .Total
Gap hitaJiSJUL^; Wts

Vent Area ; Wts

Outs. Area 64^2l2L: Wts

Noof SegH
No of Slots
K.

lin.Circ.

L22.x//^=

Section Teeth .
Volume Teeth.
Flux Density.

Loss'^g-p. cu CflZi,Total

Weight of Iron.

Mesh.

-Throw

Cond.p.Slot ^2-^,2
Total Conds 2S0^'^^

SizeofCond X

Amp. p. sq. C^i

Length in Slots.^

Length outside j£LSum

Total Length M.

Wt of i.ooo^'^ ^Total

Res. p. i.ooo2:21.Total
Watts ^JBStr±

Surface p. f^^tt^
Watts p. Sq. g^i

00/2

5-7

Afi.

/♦

J2lS

.£12.

HJOQ.

69QO

j±e£_

2LS.

^BO

TSo

ISO

Joo

nop

fB,000

2£e_

loas

iS4-0

[02 K g

iand n

1152

'0€2 sj^cm

JM.

A&ln.

2^'4L

700

'037

JX. Slots

^f'MI'^

2-8

X40

I I
I I
I I

%5> ^ T '

J9'-^

X"X

S3. Slots

no,Vcnt9

"^'5net-^ y

Field

Rotor.

Dia.

\ Total Air Gap

Gap Co-efl. K,

3tBM.

jia.

Pole PitchZ^ Pole Arc
Kf

jLZZ.

Flux per Pole.
Leakage n.l f.L

'6S

iveaiu^e n.J t.i

• Area / 7 Flu:k density

Unbalanced Pull

No. of Seg £_J Mn.Ciic.

No.of Slots -il_ x/'/«

K, Section^fi.-—

IgZ.

£±.

Weight of Iroi

A.T.p Pole n. Load'.
A.T.p.Polef.Load]
Surface

Bars \ Rings

Surface p. Watt

I*. R

I. R.

Amps.

Z7Q I 2/Q

No. of Tumsu
Mean4r9tn»

380pcnbar: ilO^ Vn ring.

IQsm

Total Length 960 crris. 24-0 1 fns^

Resistance OOISS 00017

Res. per i.ooo ^ *I7 ' '06 ■

Size of Cond \ fsq. cnt Ssg- cnrs.-

Conds. pep Slot.
Total

Length

Wt. per I.ooo.

Total Wt

Watts per Sq.-

Paths in parallel

^tiirrai 6affe

I t20 kHograms

Magnetization Curve.

Core

SUtor Teeth
Rotor Teeth
Gap

Pole Body
Yoke

•letlon. L«ngth

IL^

£.QS^

MIL

1750 OS

.Volts.

A.T.i»-

A.T.

..vr(7iC^VOlt8.

esoo 2

A.T.p-c*4a T.

laooo

12^00 JI&.

6170

100

J2^

280

JSl

870

A.T.P

Volts.

A.T.

Conr^mutator.

.Speed

Dia

Bars ..
Volts p. Bar.
Brs. p. Arm .
Size of Brs. .
Amps p. sq. .
Brush Loss .
Watts p. Sq. .

crnciENCY.

Friction and W.
Iron Loss

Ijload.

Full.

^Qtor

Arm. &c. r-'R.
Brush Loss .

:J2a.

Output
Input

Efficiency ^.

2 €7

28-6

L.UL.

Mag. Cur. /7-^ Loss Cur
Perm. Stat. Slot *

,. Rot.Sk>tJC

„ Zig-zag
2 x /^ X 7-04-
If 77 >^ 197 X /5 =^5600

6
1-6
h2^
4-0

nnd26x29 X64- =4620
197 Amps; Tot. /a^j?^
T^'09 ; X. = 4 36

^»/s» ; '. « '69 + '52

Imp. ^123-^ 192 = 4-5 ,
Sh. cir. O.r /////> A I92in/s.

Starting Torque

Max. Torque 2 6 times

Max. H.P 2-4 V'm^

Slip .^-252

Power Factor 'SS

472 DYNAMO-ELECTRIC MACHINERY

stator in star or in mesh. If the motor were to be started on a resistance in the
rotor circuit, we would prefer a star-wound stator, because the number of wires
would be fewer and the copper factor better. This motor, however, is to be of the
squirrel-cage type, and is to be started by connecting it across the mains first in
star and then in mesh. The stator must therefore be designed to work at full
voltage when mesh connected. If we take the coefficient K^ as 0-415 for a 3-phase
star-connected stator, it will be 0-24 for a mesh-connected stator. Thus we arrive
at our voltage formula (see page 25) :

490 =0-24 xl -66 xZaX 0-106,
Za = 1152,
1152 -=-72 = 16 conductors per slot.

For ease in winding, we choose round wire double-cotton covered, 0-031 sq. cm.
in cross-section. There are thus 32 wires per slot, two in parallel. The current
per double conductor is 22-6 amperes, so that the current density is 364 amperea
per sq. cm. As the coils are huddled closely together in the manner shown in Fig. 138,
this current density will be found quite high enough. The size of slot is given on
the sheet. The mouth of the slot is made 0-35 cm. in order to facilitate the intro-
duction of the wires.

Cooling conditions. The iron loss as worked out on the calculation sheet amounts
to 716 watts, and the buried copper loss 370 watts, so that the total watts to be
dissipated by the iron surfaces of the stator are 1086. From the rules given in
Chapter X., we see that for 45® C. rise we can get rid of 1340 watts, so we have not
much in hand. The rotor should be provided with a fan at each end to blow air
over the stator winding. It is a good plan to make passages behind the stator
frame as shown in Fig. 413, so that the air can get away readily and at the same
time cool the cast-iron of the frame.

The air-gap. On a small motor like this the air-gap may be made just as small
as is consistent with ensuring mechanical clearance under practical working con-
ditions. A clearance of 0 08 cm. will be found to be sufficient. With very good
workmanship and ball-bearings it could be reduced still further, but it will be
seen that the ampere-turns on the gap 0-08 cm. long only amount to about one-half
of the total ampere-turns, so that it is not worth while to reduce the length. More-
over, a reduction of the length will increase the zigzag leakage, which is already
large on those motors with a big ratio of tooth pitch to air-gap.

Itlagnetizing current. It will be seen that, owing to the short air-gap and wide
opening at the mouths of the slots, the gap coefficient Kg is fairly great, 1 -22. If
the ampere-turns in the gap are

6170 X 0 -08 X 1 -22 X 0-796 = 480,
the total ampere -turns per pole are 870.

In working out the magnetizing current we must remember that the stator is
mesh connected. The number of conductors per pole is 192, so in the mesh

0437/,«,„x 192 = 870,

/,m« = 10-3,

In, in the star = 10 -3 x 1-73 = 17-8

—

a-ri

•—

w—

»«»—

9-.

«—

«o —

*—

«o-

0-.

w —

♦ - ■

n
s

c
o

__ -:: t

r-l

474

DYNAMO-ELECTRIC MACHINERY

Squirrel cage. A fairly simple way of estimating the resistance r^ i of the squirrel
cage as measured by its effect on the stator current on short-circuit is to make a
rough estimate of the total losses in the bars and end rings with some particular
current (say full-load current) flowing in them. First find the total ampere-wires
in the stator with full-load working current. This is 20,200. Divide by 53 bars
in rotor, and we get 380 amperes (virtual) per bar. Assuming a distribution on the
rotor similar to that in the stator, we arrive at 1100 amperes (virtual) in the ring.
Now work out the total resistance of all the bars, allowing something for joints.
This is 00016 (cold) or 000185 (hot). The total loss in the bars will be

380 X 380 X 0 00185 = 270 watts.

Now, the resistance of both rings measured right around each ring sO 00017
(hot), so the loss in the rings is

1100 X 1100 X 000017 = 210 watts.
Thus, the total loss in the rotor winding at full load is 480 watts.

, <9'.— -

\S'

7'^

- -i.7- -•

0 ff N' >" ^

Fig. 415. — Circle diagram of 85 h.p. induction motor, particulars of which are given on page 46S.

Scale 1 mm. =2 amperes per phase.

Now, the resistance of the stator winding is 1 -77 ohms, so the working current
of 17 '5 amperes will cause a loss of 545 watts. Therefore

r^. 1^480
r," 545'

If the resistance of the limb of the stator be taken at 1 -77-^3 =0-59, then
fg J =0-52 ohm.

Slip. Taking the l^R loss in the rotor at 480 watts and the output of the rotor
at 25,400 watts, the input to rotor is 25,880 watts. The slip therefore will be

480
25:880' " *^°"* 2 %•

Leakage fkctor. The method of working out the stator and rotor leakage is
exactly similar to that described on pages 420 to 428. The figures are given on
page 471. The short-circuit current is 192 amperes per phase. The power factor
is found from Fig. 415.

CHAPTER XVIII.
CONTINUOUS-CURRENT GENERATORS.

Continuous-current generators of the hetero-polar type necessarily generate
alternating current within the armature itself, the current being transformed to a
continuous stream by means of the commutator. The rules, therefore, that have
been given for the calculation of the magnetic circuit, the iron losses and the copper
losses of alternating-current generators are applicable to continuous-current
generators.

There are, however, certain matters which are peculiar to the c.c. generator,
the most important of which is the bringing about of good commutation. We
propose to consider some of these in this chapter. It is not within our province
to describe the various types of winding which are used on these machines : they
are fully dealt with in many excellent text-books. We shall assume that the student
is familiar with o.c. windings ; we shall only deal with the various types of winding
in 80 far as is necessary to determine the choice between one type and another.

Leaving out of account open-circuit windings^ which are only used in very
special cases, the winding on an ordinary C.C. generator consists of a mesh-con-
nected multi-phase winding, the number of phases being equal to the number of
commutator bars per pole. By making the phases very numerous, and by making
a connection between each phase and a commutator bar, we are able to generate
a very uniform voltage. This voltage is made up of the sum of the £.m.f.'s gene-
rated in all the phases, except the one or two which are short circuited by the
brushes, and are undergoing commutation. Two advantages accrue from the
increase in the number of phases : in the first place, the variation of voltage caused
by the cutting-in and cutting-out of a new phase is reduced ; in the second place, the
self-induction of each coil under commutation is reduced. If we can increase
the number of phases until each consists of only one turn, we have approached the
ideal condition. In some cases it is even advisable to go further than this, and
make a connection to a commutator bar after each half-turn. So important is
it to keep down the value of the b.m.f. necessary to reverse the current in the
coil under commutation, that in large machines a number of circuits are arranged
in parallel, there being as many poles provided on the field-magnets as there are
circuits in parallel : thus only a fraction of the current to be delivered is dealt with
by one pair of poles.

476

DYNAMO-ELECTRIC MACHINERY

Crommutation. The conditions which are to be met in order to secure good
commutation may be shortly stated as follows : We shall speak of one section of
the winding, the ends of which are connected to successive commutator bars and
which really constitute one phase of our multi-phase generator, as an armature
coil. In large machines each coil consists of only one turn. In Fig. 420 are shown
diagranmiatically a number of coils connected to conmiutator bars : let us fix our

+ /.

dmrnxL-

.2^

Fio. 420. — Gommatator bars passing under brush.

attention upon one ot these connected between bars 2-3 as the armature moves
forward in the direction indicated by the arrow. We see that the current through
the coil is going from left to right before the coil reaches the brush, and going from
right to left after the coil passes the brush. The interval of time during which

the coil is short circuited by the brush is <c = — > where 6^ is the breadth of the

brush in centimetres, and Vc is the velocity of the commutator surface in centimetres

0\ >

\ »

FIO. 421.

per second. During this interval of time the current must be completely reversed \
and if we are to have no tendency to spark when the bar 2 leaves the toe of the
brush, the current in the coil 2-3 must have grown to be exactly equal to the current
in coil 1-2. This reversal of the current is most satisfactorily brought about by
introducing an E.M.F. (called here the commutating e.m.f.) into the coil during the
interval of short circuit. If this commutating e.m.f. is too small, it will fail to build
up the reverse current to the value of the normal armature current before bar 2

CONTINUOUS-CURRENT GENERATORS 477

leaves the toe of the brush, and we must rely upon the resistance of the carbon
brush to force the commutation just at the last instant (see curve u in Fig. 421);
while, on the other hand, if it is too great, it will build up the reverse current to
a value higher than the normal armature current and cause sparking from '' over-
commutation," if the resistance of the brush is not sufficient to force the current
to the right value at the last instant. It will be seen that the ideal commutating

^■'a

-4

Fig. 422.

E.M.F. is one which will be zero at no load, and which will increase as the load comes

on, so as to be equal at all times to -~xL, where L is the coefficient of self-induction

of the coil under commutation and - — is the rate of change of the current which is

necessary to produce an exact reversal in the interval of time tg. As L is fairly
uniform for a wide range of la, it is desirable in general that the commutating
E.H.F. shall increase in proportion to the load. This commutating b.m.f. will
be generated in the coil under commutation if during the commutation interval the
coil is moving under a commutating pole, as illustrated in Fig. 429.

+ 4

-'a

no. 423.

If the commutating e.m.f. is nearly constant during the commutating interval,
the rate of change of the current will be almost constant, and the current in the
coil will change from positive to negative in the manner shown by the full line in
Fig. 421. If we bring about conmiutation according to this straight line law, the
conmiutating e.m.f. will be a minimum, but it is safer to make the rate of change
greater in the middle of the commutating interval and smaller at the end of the
interval. The resistance of the brush considerably afPects the shape of the curve
and helps the current in the coil to reach the right value before the brush leaves the
bar. Thus, if the commutating e.m.f. is too small, the resistance of the brush hurries
up the commutation towards the end of the interval (see curve u in Fig. 421), while
if the E.M.F. is too great, the resistance prevents excessive over-commutation and
forces the reversed current to come down to the normal value just before the brush
leaves the bar (see curve o in Fig. 421). The effect of the brush is to give something
in the nature of a back e.m.f., which opposes the flow of current out of the leaving
bar (assuming the bar is positive). The back e.m.f. which the brush exerts is only

478

DYNAMO-ELECTRIC MACHINERY

of small value (some two or three volts), so that if the adjustment of the com-
mutating pole is so imperfect as to call for a greater correcting influence than can
be exerted by the brush, sparking will occur. It is not well to rely on the brush to
correct an error in the adjustment of the commutating voltage of more than 1 or
1 -5 volts. For this reason one tries to keep below 10 volts the commutating voltage
required to reverse the current in an armature coil, so that if there be a 10 per
cent, error in the adjustment of the pole, the carbon brush can prevent sparking.
If we have a commutating voltage of 20 volts, an error of 10 per cent, would require
the brush to exert a back correcting pressure of 2 volts, and the commutation might
not be satisfactory. Certain kinds of brushes are capable of giving up to 3 volts
correcting e.m.f. before showing much sign of distress.

Fio. 424. — Showing p; the pitch of the slots ; bpf the breadth of the brush referred to the
periphery ; Cp, the breadth of a commutator bar referred to the periphery.

The main difficulty in providing this commutating pole and exciting it by the
right amount of ampere-turns arises from the fact that the armature is itself a power-
ful electro-magnet, one of whose poles is directly on the place where we wish to fix the
commutating pole, and the direction of the magnetization of the armature tends to
produce a flux of the opposite kind to that required for commutation. If, there-
fore, we put an iron pole-piece in the place where the commutating pole is desired,
the armature magnetomotive force tends to produce a flux through that pole which
is of the wrong polarity ; and before we can begin to get a flux of the right polarity
we must put a number of ampere-turns on the commutating pole, which is greater
than the number of ampere- turns operating in the armature. Having balanced
the armature magnetomotive force, we must put on enough additional ampere-
turns to produce a flux of the right value to bring about good commutation. Thus
we have on the armature and commutating pole two powerful magnetomotive

CONTINUOUS-CURRENT GENERATORS

479

forces which oppose one another ; and it is only the difference between them which
is effective in producing the conxmutating field. Thus it comes about that there
is often a very heavy magnetic leakage from the conxmutating pole to the adjacent
main pole, and this magnetic leakage tends to bring about the saturation of the
iron on the conxmutating pole at heavy loads, and destroy the correct proportion-
ality of the commutating flux. For this reason the cross-section of the iron on
the commutating pole, particularly near the root, should be fairly heavy. A
calculation should be made of the magnetic leakage, and sufficient iron provided,
so as to avoid saturation. This is of particular importance on commutating poles,
because the flux is produced by the difference between two magnetomotive forces ;
and as the difference is not very great compared with either one of them, a very
Uttle reluctance added to the magnetic circuit destroys the proportionality of the
flux. The number of effective ampere-turns to be put upon the commutating pole

TZIl

- - v<

FIO. 426.

can be arrived at by the following considerations : Take all the conductors in
any one slot, and see what change of total current occurs in those conductors as
the slot moves past the brush. With a full-pitch winding the whole of the current
in the slot will have been reversed during the interval in which the commutator
bars to which the conductors are connected pass under the brush. If ps is the
pitch of slots in centimetres, and bp the breadth of the brush increased in the ratio

of j^, and c« the width of the commutator bar increased in the ratio ^. where

dc ^ dc

de is the diameter of the commutator and da is the diameter of the armature, then
the interval of time taken for the current in the slot of an armature with a full-pitch

winding to completely reverse is — — ^ — -^, where Va is the peripheral speed of the

armature (see Fig. 424). In considering the leakage flux per centimetre length of iron
around the conductors lying in one slot of a full-pitch winding per ampere passing^
we shall make use of various dimensions indicated in Figs. 425 and 426.

480 DYNAMO-ELECTRIC MACHINERY

We shall use the following symbols for the magnetic flux per centimetre of axial
length of iron per ampere in the slot :

Ln = the effective flux crossing the body of the slot.
Ljc = the flux bridging from the tops of the teeth along the air-gap.
Le=the flux bridging across to the commutating pole and back again.
Lg = the flux encircling the end connections of the armature coil.

In the case of a wire-wound armature coil we may take

^-'■»'(I4:)-

In the case of strap coils, the magnetic flux through the strap cannot change
quickly on account of the eddy current generated in the copper. In this case, as
a rough approximation, we may take

^-'•»'(^/5.>

We have i:i=0.921ogio(j*),

and L,=0-46|(logioj^-0-2).

Where there is a commutating pole, it is usual to neglect the term Lt, because
Lc takes its place, except in those cases where the length of air-gap under the com-
mutating pole is very great.

Let us denote the sum of the leakage fluxes per centimetre length of iron by Lf.
Then Lt = Ln + Lc + L,

in the ordinary commutating-pole machine.

Now, if the axial length of the commutating pole is the same as the axial length
of the iron, the flux-density Be in the gap necessary to bring about good commuta-
tion * will be nor amperes per slot

•

♦ The proof of this formula is as follows. Assume that Be is constant over a part of the
periphery of the armature, having a width p,+bp- Cp, The conductors moving under the
commutating pole, by reason of their motion, cut the flux from the pole Be x (p« + 6p - Cp) x 2^,
while the current in one slot is being reversed. During the same period the flux created by the
current in the slot changes from positive to negative, so that the cutting due to this latter cause
is 2 X Xt X 2i X amperes per slot. In order that these two effects shall neutralize each other

Be X (p, + 6p - Cp) = 2Lt X amperes per slot.
The following references to papers on the theory of the commutating pole will be of service :
Arnold and La Cour, "Commutation," Trans. Intemai, Elec, Cong,, p. 801, 1904; see
also Worrall, J(mm. I.E.E., vol. 45, p. 480 ; Page and Hiss, ibid., vol. 39, p. 670; " Improved
Arrangement of Commutating Poles for Dynamos," Engineer, 106, p. 181, 1908 ; " Motors
with Commutation Poles," W. Siebert, Elektrot. Zeitschr., 30, p. 466, 1909 ; " Interpole Designs,"
W. B. Hird, I,E,E. Joum., 43, p. 609, 1909 ; " Auxiliary Poles for Direct-current Machines,"
J. N. Dodd, Amer. I,E.E., Proc. 28, p. 467, 1909 ; " Reactive Effect of Auxiliary Poles in d.o.
Machmes," F. Punga, Elek. u. Masckinenbau, 29, p. 306, 1911 ; "Design of Auxiliary Poles,"
A. Brunt, Elec. Rev. and West. Electr., 59, p. 607, 1911 ; "Hunting of Direct-current Interpole
Motors," E. Rosenberg, Electrician, 67, p. 670, 1911; "Calculation and Experimental Determi-
nation of Mean Reactance Voltage," J. Liska, Ehk. u. Masckinenbau, 30, p. 826, 1912 ; " Leakage
Coefficients of Commutating Poles," L. A. Doggett, Electrician, 69, p. 821, 1912 ; " Calculation
of Interpoles," De Bast, Assoc. Ing. El. Li^ge, Bull. 13, p. 208, 1913 ; " Armature Reaction and
Characteristic Curves of d.c. Dynamos," Guilbert, Lumtere Elec., 22, p. 69, 1913.

CONTINUOUS-CURRENT GENERATORS 481

where p^, hp and Cp are measured in centimetres, and have the signification shown
in Fig. 424. This is on the assumption that we wish to bring about commutation
according to a straight-line law. If we wish the commutating curve to follow
approximately a sine curve, then we will have to shape the commutating pole so
as to give a fringing flux, and the maximum value of Be will be obtained from the

expression b. = 2 -8 x Z, x 'amperes per slot

P« + Op - Cp

A somewhat similar eHect can be obtained by making the throw of the armature
coil one slot smaller than a full pitch, and making the width of the commutating
pole just about one slot pitch. We then get a commutation curve like that shown
in Fig. 423.

If the axial length of the commutating pole shoe, U, is less than the axial
length of the armature iron, fe, then two corrections are necessary : in the first
place Lc must be reduced in the ratio Ic/Uy and in the second place we must
increase B<. by the ratio Ullc-

If the armature has not a full-pitch winding, the coil undergoing commutation
under a N. pole will not lie in the same Blots as the coil undergoing commutation
under a S. pole. The interval of time during which the leakage flux across a slot
is reversed will be twice as long. The effect is approximately the same as if the
flux leaking across the slot were reduced to one half, so for short throw coils we divide
Ln and L^ by 2. Le and L« are, however, practically imaffected.

Magnetic OBcillationB. Besides the e.m.f. produced in the armature con-
ductors by the reversal of the leakage flux above-mentioned, there are other
E.M.F.'s which must be guarded against. If the pole is not bevelled and the slots
per pole are few, the swinging of the flux (considered on page 313) and the move-
ment of the conductors under the pole produce ripples in the continuous voltage
generated between a pair of brushes. The current supplied by the generator will
also have ripples in it, and these set up high frequency e.m.f.'s in the conductors
under conmiutation, and may cause sparking troubles. For this reason it is not
well to have too few slots per pole. The fringing flux from the comer of the pole
may cause trouble in the conductors under commutation if it swings about and
generates e.m.f.'s which are not proportional to the velocity of the conductors.
This trouble is obviated by bevelling the pole and by arranging to have at least
three or four slots between the main poles. Flux swinging may occur under the
commutating poles themselves and produce alternating £.m.f.'s which are super-
imposed upon the legitimate e.m.f. of commutation. The greater the number of
slots in the commutating zone and the greater the air-gap under the commutating
pole, the less these effects will be. The most perfect way of getting rid of these
effects is to mount the punchings of the armature so that the slots are not parallel
to the axis of the generator, the position of a slot at one end of the armature being
exactly one jslot pitch further ahead than the position of the slot at the other end
of the armature, as shown in Fig. 533. This skew mounting of the punchings
may be either on the rotor or on the stator. Sometimes the edge of the pole, instead
of being parallel to the axis of the machine, is given an inclination to the axis,
80 that one comer is one tooth pitch further ahead than the corner at the other

W.M. 2 H

482 DYNAMO-ELECTRIC MACHINERY

end of the armature. Upon the whole, the bevelling of the pole is the cheapest
method, and has the additional advantage that it diminishes the noise caused by the
rotation of the armature at the same time as it diminishes the magnetic effect which
we have been describing.

Brush gear. The problem of how to collect an electric current from a quickly-
moving conducting surface has been a very difficult one, and it has only been
partially solved. In the early days of the d3aiamo, nothing seemed easier than to
use a metal wire brush on a copper commutator. But the wear between metal
and metal is too great in these days, when machines are expected to yield thousands
of amperes day after day with little or no attention.

The use of a carbon brush on a copper conamutator gives us very little wear
when the materials are of the right quaUty. Under good conditions, a carbon
brush takes a highly-polished surface, which makes no impression mechanically
on the tough copper of the commutator.

If no current passed between the commutator and brushes, most commutators
would run for years without showing any appreciable wear. It is not mechanical
wear that gives trouble. The main difficulty arises in keeping the working face of
the brush in close contact with the commutator. If there is any distance at all
between the copper and the carbon, the current can only pass from one to the other
by means of a short arc. It is this arc that causes nine-tenths of all the commutator
troubles brought to the notice of the designers of continuous-current machines.
If this arc is extremely short (less than one ten-thousandth of an inch) it may not
do more than provide the requisite voltage drop between carbon and copper. When
it assumes a length of one or two ten-thousandths, it begins to be troublesome,
and may have the effect, when the current is passing from copper to carbon, of taking
copper off the commutator and plating it on to the brushes. Thus, if some of the
bars of a commutator are just a little lower than others, so that the carbon brushes
do not quite touch them, there will be a tendency for capper to come off the low
bars and make them lower still ; and after a few weeks' running a '' flat " develops
on the place where previously the lowness of the bars could not have been detected —
except, perhaps, by a slight difference in the colour.

The necessity of keeping the carbon brushes in close contact with the com-
mutator, makes the design of the brush holder a matter of the greatest importance.
It is not within the province of this book to treat at length upon the mechanical
design of brush holders — that subject would require a book to itself. All that can be
done here is to point out the main features that a good brush holder should possess.

A brush holder should be firmly supported. Not only should the arms which
support the holders be stiff, but the part of the machine to which they are attached
must be very rigid. The construction of the box of the holder, or the part which
holds the carbon, should be sufficiently rigid to resist any distorting forces that come
upon it during the running of the machine.

While the carbon must be free to follow any slight eccentricity in the commutator,
it must be held so that it does not tilt or change the angle of the face presented to
the commutator. For this reason the " box type " holder, in which the carbon
can slide parallel to itself, has found more favour than the pivotted holder, in which
the carbon tilts through a slight angle when the commutator nms out of truth.

CONnNUOUS-CURRENT GENERATORS

483

Fig. 427. — Giving the approximate voltage drop at bnuhes
under good conditions with ordinary carbon brushes.

Box holders have sometimes the drawback that they do not fit the brushes well

enough. Either the brush is so tight that it cannot slide, or it is so loose that it

shakes about. For this reason, it is best to supply the box with a side spring that

keeps the brush pressed lightly but definitely against the side of the box against

which it is intended to slide. •

The holder should be provided with a spring for pressing the carbon against the

commutator ; and this spring should be capable of easy adjustment while the machine

is running. A pressure of 1 lb. to

IJ lbs. per square inch of contact

area is generally sufficient. In cases

where the commutator can be made

to run very true, smaller pressures

can be used successfully.

The brush holder should be so

made that the brushes can be taken

out and inspected readily without

altering the adjustment of the spring

tension. Nearly all brushes are now

supplied with flexible leads to carry

the current from the brush. These

flexibles should be made very ample, because it is found in practice that a brush

is often called upon to carry much more than its share of the load.

Besistance of the brushes. The voltage drop which occurs at the contact sur-
face of the brushes when current is passing depends upon (a) the kind of brushes, (6)

whether the brush is positive or negative, (c) the current density, {d) the mechanical

pressure employed, and (e) the state of the commutator.

With brushes of ordinary hard carbon, the voltage from copper to carbon and

from carbon to copper varies with the current density,* as shown in Fig. 427. These

results were obtained on a com-
mutator which had been a long
time in service, so that it had
acquired a fine polish and was
completely free from low bars or
high mica. The pressure employed
was 1^ lbs. per sq. inch. Where a
higher mechanical pressure is em-
ployed, the voltage drop may be
lower. Fig. 428 shows how the
voltage drop changes with the

mechanical pressure. In the experiments from the results of which these curves

were plotted the contact conditions were not as good as in the experiments recorded

in Fig. 427. The voltage drop remains practically constant for conamutator speeds

between 2000 and 5000 feet per minute.

Some soft graphite brushes are specially designed to give a low contact drop.

For instance, the "l.f.c. 2" brush of the Le Carbone Company, when worked

* Arnold and La Cour, Trans. Intemat. Elec, Cong,, 1904, p. 801.

Fio. 428. — Giving the approximate voltage drop at brushes

with oifTerent pressures.

484 DYNAMO-ELECTRIC MACHINERY

at a cuirent denBity of 55 amperes per sq. in. and at a mechanical pressure of If lbs.
per sq. in., gives a total drop of only 1 -2 volts on positive and negative brushes
taken together. The Morgan Crucible Co. also make graphitic brushes giving a
very low contact drop. Where the commutator runs very true and where the
mechanical conditions are exceedingly good (as, for instance, of the radial com-
mutator illustrated in Fig. 437), voltage drops as low as 04 have been obtained at
pressures not exceeding 1^ lbs. per sq. inch on graphitic brushes.

Where a certain amoimt of copper is added to the carbon, the voltage drop is
reduced, but the addition of this copper always tends to increase the coefficient of
friction between the brush and the commutator, and when enough copper is added
to substantially lower the brush drop, it is found that the wear on the commutator
is very much increased.

The coefficient of fiiction of a hard carbon brush working on a commutator in
perfect condition will vary from 0-3 for a speed of 500 feet per minute to 0-25 for
a speed of 5000 feet per minute. For graphitic brushes the coefficient of friction
varies from 0-27 for a speed of 500 feet per minute to 0 15 for a speed of 5000 feet
per minute. The coefficient of friction, however, is a very uncertain quantity,
and varies very greatly for slight differences in the state of the commutator surface,
and with the type of brush-holder employed. In calculating the losses due to
friction, it is well to take the coefficient at 0*25 for graphitic brushes and 0-3 for
hard carbon.

Oharacteristic curves. A great deal has been written on this subject in books
dealing with the theory of continuous-current machines, so that it is unnecessary
to discuss it here. Where a machine is compound wound the manner of deter-
mining the number of turns of series winding is in practice very simple. A
magnetization curve (see pages 281 and 489) is plotted, and the number of ampere-
turns required to give the increased voltage on full load ascertained, allowance
being made for cross-magnetization and voltage drop in the armature and brushes.
The shunt ampere-turns at full voltage are deducted from these, and the remainders
give the required series turns. In practice one puts on a few ampere-turns in
excess, because it is so easy to adjust the machine when it comes on t€st by
means of a diverter.

THE SPECIFICATION OF CONTINUOUS-CURRENT GENERATORS.

Regulation. There are not so many qualities to take into accoimt on a con-
tinuous-current generator as in the case of some of the machines dealt with in other
chapters ; and the specification can therefore be made very short and simple.
One feature which the purchaser or his adviser must look to is the regulation-
characteristic. This will depend on the nature of the service for which the generator
is intended. For traction work it is usual to install compoimd-wound generators,
and in the past 10 per cent, over-compounding has commonly been asked for.
This is in general more than sufficient to compensate for the drop in feeders. The
purchaser should make a rough estimate of how much over-compounding will be
required to make Ihe operation satisfactory in practice, and not call for more than
he requires.

CONTINUOUS-CURRENT GENERATORS 486

When calling for compound-wound generators, the specification should state
whether the series winding is to be connected on the positive or negative side of
the machine. Where generators are already installed with which the new machine
must run in parallel, it is well to state the voltage which at present exists between
the equalizer bar and the positive bus-bar (if the series winding is on the positive
side) at full load on the power-house. This will enable the designer of the new
machine to adapt the resistance of his series winding in the most economical way.
It is also well to state what the actual rise of voltage is between no load and full
load. The designer of the new machine, having found out the change in speed which
will occur with the prime mover, can arrange his winding and the characteristics
of his machine to meet the conditions.

For ordinary town lighting and power supply, it is usual to employ shunt
machines, either hand regulated or controlled by an automatig regulator.

Where shunt machines intended for important work are specified, it is well to
give the characteristics of the generators at present installed and the actual drop
in voltage when full load is thrown on and after the prime mover has settled down
to its full-load speed.

If the duty for which the generator is required is specified, it will sometimes
assist the manufacturer to adapt his machine more exactly to meet the required
conditions.

The first specification which we give as an example relates to a small -generator.
For this the specification should be as simple as possible, and should not contain
any clause which will prevent a manufacturer quoting on his standard machine,
otherwise the prices quoted will probably be higher than lowest competitive prices.
The requirements in performance only should be stated. It is not wise to call for
a certain efficiency : it is better to ask the Contractor what the full-load losses on
his machine are, and then, on comparing tenders, an allowance can be made in
prices on the basis of the losses.

486

DYNAMO-ELECTRIC MACHINERY

SPECIFICATION No. 10.

75 K.W. CONTINUOUS-CURRENT BELT-DRIVEN GENERAiTOR
Clauses 1, p. 269 ; 21, p. 333; or 170, p. 519.

Characteristics 150. There shall be supplied a shunt- wound continuous-
current generator having the characteristics set out below :

Pulley.

Delivery.

Statement of
Losses.

Normal output

Normal voltage at ter-
minals

Voltage adjustment on
rheostat

Normal current

Speed

How driven

Size of steel pulley to

76 K.w.

625.

500 to 630.
143 amperes.
760 R.P.M.
Belted.

12 in. dia., 10 ins. wide.

be suppUed
Temperature rise after
2 hours full load run 46° C. by thermometer.

56° C. by resistance.
Over load 25 per cent, for 30 minutes.

Temperature rise after
30 minutes over load 60° C. by thermometer.

70° C. by resistance.
Puncture test 1500 volts (alternating) apphed

for 1 minute between wind-
ings and frame.

161. The generator shall be provided with a pulley of the
size above specified and be mounted on slide rail with belt-
tightening screws.

162. The contract includes the delivery of the generator at
the purchaser's works in , but does not include
erection or starting-up.

163. The contractor shall state the amount of the follow-
ing losses in the generator which he supphes :

1. Bearing friction and windage losses (at no load).

2. Iron losses (at no load).

3. Armature and field copper losses at full load,
allowing for temperature rise.

CONTINUOUS-CURRENT GENERATORS 487

164. The generator shall be run for two hours at full load Tests.
at the contractor's works in the presence of the purchaser's
engineer, without showing any sparking at full load, and with-
out injurious sparking on over load. At the time of this run
tests shall be made to see if the machine has the characteristics
set out in Clause 160, and measurements shall be made of
the losses above specified. If it is found to comply with all
the conditions it shall be accepted without further tests. But
if during the first six months after deUvery any defects in
the construction or performance become manifest, the same
shall be immediately rectified by the contractor at his expense.
Any time between the reporting of defects and the remedying
of same shall not be counted in the six months' period of
maintenance.

DESIGN OF A 75 K.W. CONTINUOUS-CURRENT GENERATOR.

525 volts ; 144 amperes ; 750 R.P.M.

Small belted motors and generators are now usually built with a fan at one end
of the armature, of the kind illustrated in Fig. 429. By this means such very good
ventilation is ensured that fairly large outputs can be obtained from small frames ;
in fact, for a machine not larger than 75 K.W., it is possible to take a D^l constant
of 4 X 10* cu. cms. In fixing diameter and length, we must remember that the
machine will (if it is to be economically manufactured) constitute one of a
line of generators whose output may vary from 1 K,w. to 100 K.w. ; and in such
a line it is usual to build several machines of different outputs on the same
diameter, the length of iron being changeable so as to make an economical machine
for each output. It thus comes about that any given machine in the standard line
may either be of moderately large diameter and short length, or of smaller diameter
and greater length, according to the accident which puts it upon one frame rather
than the frame smaller. A considerable difference of opinion still exists between
designers as to how far it is economical to lengthen the core of a given frame parallel
to the shaft before going up to the next frame with a shorter core. It appears,
however, that if we take into account the cost of material and labour on machines
manufactured in large numbers, the most economical machine will be one which
has a ratio of length of iron to pole pitch lying between the limits 0*5 and 0-8 ; and
without going into the labour and material costs in great detail in any given factory,
it would be difficult to state more exactly the best possible dimensions. In fact,
even if the ratio of length of iron to pole pitch lies considerably outside the above-
mentioned limits, it does not follow that the cost per kilowatt will be very much
increased. The best ratio of length of iron to pole pitch is largely dependent upon
the question whether the designer chooses to build a " copper " machine — ^that is
to say, one in which the laZa is great — or an " iron " machine — one in which the
AgB is great (see page 8).

488 DYNAMO-ELECTRIC MACHINERY

This ratio of length of armature iron to pole pitch is also controlled by the shape
of the field pole. If we decide to use round poles, we cannot well, on a four-pole
machine, make the ratio much greater than 0-8. It will, however, be possible with
round poles to have three standard sizes of frame, using different diameters of pole
body : three convenient ratios in this case are 0-5, 0-65 and 0-8.

For the 75 K.w. machine under consideration, we will adopt the ratio 0-66,
making the diameter of the armature 43*5 cms. and the length 22 cms.

A calculation sheet is given on page 489. We begin by filling in the main data
of the machine.

The best number of poles to take when designing a c.c. generator is controlled
partly by the considerations given on page 10, and partly by consideration of
the amperes to be collected at each brush arm. Where a generator is of small
output and the total current is not great, say under 800 amperes, there is no advan-
tage to be gained in making more than four poles. An increase in the number
of poles increases the cost of labour. Moreover, with small machines, it would
be difficult to get in enough commutator bars per pole if there were six poles. Four-
pole machines are cheaper than two-pole machines, foi*the reasons given on page 12,
except for very small sizes, where the labour is the main consideration. Where
the current output is very great, the number of poles is increased so that the current
per brush arm may not be excessive. Five hundred amperes per brush arm can be
dealt with very satisfactorily with properly designed commutating poles.

Other considerations controlling the number of brush arms are given on page 567.

We will therefore choose four poles for this machine and proceed to fill up the
calculation sheet.

To obtain 525 terminal volts at full load, we should allow for the generation of
540 volts at no load. The amperes per terminal will be 144, the cycles per second
25, the speed 750 R.P.M., or 12-5 revs, per sec. With a two-circuit winding, the
amperes per conductor will be 72 ; and with 4 brush arms the amperes per brush
arm will be 72. The specified temperature rise is 45° C, and the over-load capacity
20 per cent, for one hour.

Type of winding. To generate 525 volts on a machine as small as 75 K.w., we
should of course have to employ a two-circuit winding. The considerations which
determine the choice of the kind of winding are given on p. 511. The method of
settling approximately the number of conductors might be as foUows :

Magnetic loading. The circumference of the armature is 137 cms., and the
area of the working face, Ag^ 3020 sq. cms. (see calculation sheet, page 489). If we
work with a maximum flux-density of 8500, our AgB will be 0-256 x 10® ; and this
will require, at a speed of 12-5 revs, per sec, about 250 conductors in series to give
540 volts (15 volts margin being allowed for drop in brushes and windings). The
exact number of conductors to choose would depend upon the number of slots that
we have in our standard punching. The number of slots in a standard punching
will by preference be an odd number, so as to enable a series winding to be con-
structed without any idle coils ; and by preference, for a four-pole machine, it will
be a multiple of 4, plus 1. The number of slots per pole ought not to be less than
9 on a machine of this size, and about 10 slots per pole would be good practice.
We will take 41 slots in all ; 41 multiplied by 12 gives us 492, which divided by

CONTINUOUS-CURRENT GENERATORS

489

l>aLtt.$.tJ9n, 19/4 Type .. ..

KW...75..... ; P.F ; Ph««e

H.P. Amps pw coad. 7«....

.CCCEN

Amps p. br. MX1D..7J2..

Amps per ter /4.4..
.. .Temp, rbe .4:S ...

.^ Poles Blec Spec ../(!?

.; Cfdta.JZ^. . i R.P.M..750...i Rotor Amps ;

RepiUtion. /^ 7f9^r<^/POrtrlOBd 20,%>JsLhour

Cnstocner : Order No ; Qupt. No ; Perf Spec ; Fly-wheel effect

Fruie#j|^0. ^. /o'y ^ - S090 ?«••• Ag B ^....; pose. laZa . . .

K. r.(6fl5. ; ... JW . Voits=:tfa5« J2:S. * J»iS « .:.Z5.6

: Circuin.

:255

D' I. V R P M

K.V.A.

.4/x/d?^

; Ann. A.T. p, pole....4r4J?<? Max. Fid. AT.

Armature. Rev.

o
o

Dia. Outs

Dia. Ins

Gross Length
Air Vents —

Mean

Opening Min

Aig- Velocity

Net Length^fiii x-Sg

Depth b. Slots

Section l^ Vol.

Flux Density.

I/xW<^^.cu.lg!^!Z-Total
Buried Cu.^g<^ Total
Gap Area ^O^O ; Wts
Vent Area_si^^^: Wts
Out^. Area ^^/2^.: Wts

c
O

■o

c
o
u

No of Segs
No of Slots
K. 2/ ■

oin.Circ

Section Teeth

Volume Teeth

Flux Density Afip

Loss;^5-p. cu C5L-Total

Weight of Iron.

/J2

Star or Mesh Throw

Cond. p Slot

Total Conds 2^6-l£LSA'2.^9.
Size of ZovA2(a5. x2il 2J<:0£^
Amp. p. sq _- 1 4^->?

Length in Slots. -:^^- _
Length outside ^4 Sum
Total Length

Wt. of i!8oo./^/ Total

Res. p. 1.000/ Total

Watts p.-ffi.

Surface p. IIL

Watts p. Sq._C5L.

IS*

Sq.J

'O012

^35

/6'S

220

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JB&HS

350

in 5

^9-2

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IS20

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lAZMI^^

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176.

374 m

960

:P2JL

JQ1C_

5S'6^fHo^rs.

11^ \09f-

surs

K ^

<
I

I
I

I I
I I

A i I

TTTT

■JT

"y^

; . 4L Slots

I
I

c— 2

>5 <t
Pole20cms,dia.* '

Field Stat

Bore

I Total Air Gap

Gap Co-eff. Kg

Pole Pitch3f-4Pole
K,

Arc

Flux per Pole^^^i<:(^

«•

ux density
Pull L

No. of Seg.l Mn.Circ

No.of Slots'
Vents __'.
K.

X =

^4-

'2S_

1*26

24" 5

686

5'/4-x/c'^

iMoa

.Section

Weight of Iron

A.T.p Pole n. Load
A.T.p.Polef.Load

Surface

Surface p. Watt-
le R

LR;

Amps.

Shunt.

384i
6200

/good

14.

No. of Turns-.
Mean 1. Turn

22Q..

4-63

i'55

340O

Oemm.

SQOO

^2QQ

j2fA

2-4

/*4.

4-1

Total Length ^IJ^QQjri.

,_,

4.6

'Si^V!AJMv::it^SQMjt^6Qii[h<^f}S^

Res. per I. (

Size of Cond

Conds. per Slot.
Total

,^4-4

iQOJ^ cms. 22^^

'Length

Wt. per i.ooo-
Total Wt

Watts per Sq

Star or Mesh

Paths in parallel

\.JZ6m
QW6_

.^&JL

6^ fd'fbgs

072

'22

679

Magnetization Curve

Core

Stator Teeth
Rotor Teeth
Gap

Pole Body
Voice

Section.

~T6S

LL9SI

3B2a

3mL

192

L«ncth

'25

18
~^2

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f2j8QQ 6

A.T.P. c:

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6260

A.T.

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iS,00d S

2090
455_
336

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.^i^O. Volts.

l^Z£l^

(MQO

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36 O 1400
2160
620

J±0_

.9:Ji.

A670

;5€<?.. Volts.

I370C 12

22^206 600
.^StQC

(7^000 4 6
13.900 //"

A.T.pr.

A.T.

120

?Pog_

2230

46^
6397

Commutator.

157h.p

D'ul.jSS. Speed

Bars 123

Volts p. Bar— Z2.
Brs. p. Arm 2

Size of Brs. 2x4-6

Amps p sq. /tf^ ^

Brush Loss ^300 fYafts
Watts p. Sq. :j1

EFFICICNCV

Friction and W —

Iron Loss

Field Loss

Arm* &c. I R

Brush Loss

IJ load.

I i'O

'92

186

4'L_
•38

JFuU.
I'O

'82

2-76

Output

I Input — --_-.
Efficiency ^>-

7'25\6'77
\ 94 \ 75

101 25 80-8

92 7 92 a

^'Q8
'8

J
I'O

'23

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'78
65

4 '41

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3 44
36 '5

'76

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•07

2-82

80j4
'927,

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91 'I

21 8

Mag. Cur. Loss Cur.

Perm. «4at. Slot 167

., ikr&rek9txfnd= 1-32

2 X

177

End

Ns.

Zig-zag
y

X
X X

Amps . Tot
.X. =

1-47

4-36

= +

Imp. V +■

Sh. cir. Cur —

Starting Torque
Max Torque _

Max. H.P

SUp

Power Factor

490 DYNAMO-ELECTRIC MACHINERY

2 gives us 246 conductors in series, a number sufficiently near to the preliminary
number 250. It is necessary to know the pole arc before we can arrive at the con-
stant Ke (see page 13). It will be seen from the drawing of the machine (Fig. 429)
that we cannot well make the pole arc wider than 24-5 cms. This pole arc with
the slight bevel shown gives us K/=Ke = 0-686 (see page 23). We can therefore
write down the formula for the voltage :

540 volts X 10«=0-685 x 12-5 x 246 x AgB.

This gives us AgB =0-256 x lO®.

If we allow three ventilating ducts each 0-75 cm. wide, we get a net length of
iron, after allowing for paper insulation, of 17 4 cms.

The size of the slots will generally be fixed by our standard dies ; and we should
as a rule have to contrive, by using several conductors in parallel, to obtain the
cross-section required in the space at our disposal. For this purpose conductors
of rectangular cross-section are very much more convenient than round conductors ;
and we have seen on page 151 that cotton-covered rectangular conductors can be
used and safely shaped into armature coils if coils of the right type are employed.
If, for instance, our standard slot has a depth of 4 cms. and a width of 1 -2 cms., we
can employ two conductors in parallel, each 0-25 x 0-35 cms., arranged one above the
other as shown in Fig. 159, to constitute a conductor having a cross-section of
0 17 sq. cm. There will thus be six double conductors per complete coil. A
rectangular conductor 0-25 x 0-35 when double-cotton covered will measure
0-28 X 0-38. Three of these side by side will take up 0-84 cm., leaving 0-36 for in-
sulation and internal roughness of slot. This room will permit of three turns of
paper and mica, together with one layer of linen tape to hold it in position. The
room in depth of the slot will be found to be amply sufficient, and where too ample
can be made up by means of strips of press-spahn inserted either in the bottom of
the slot or between the limbs of the coils.

As the current per conductor is 72 amperes, the current density will be
72 -7- 0-17 = 425 amperes per sq. cm. This we know from experience is not too high
a current density in an* armature of this type ; but strictly we ought to work out
the cooling conditions for the slot as indicated on page 224.

It next remains to work out the flux-density in the teeth. The depth of tooth
being 4, we take the diameter of the mean circle as 37 -5, giving us a mean circum-
ference of 117-5 cms. As there are 41 slots each 1-2 cms. wide, we subtract 49*2
and obtain 68-3 cms. as the total width of all the teeth. Multiplying this by the
net length, 17 -4, we obtain 1190 sq. cms. as the total section of all the teeth. Divid-
ing this into 0-256 x 10®, we obtain 21,500 as the apparent flux-density in the
tooth at one-third of its length from the root. As the ratio ^,=2-1 (see page 71),
the actual flux-density in the teeth is 21,000. From Fig. 29 (page 51) we find the
loss at 25 cycles (B = 21,000) is 0-08 watt per cu. cm. As the volume of the teeth
is 4750 cu. cms., the loss in the teeth is 380 watts. A flux-density of 21,000 is not
excessive for 25 cycles, and therefore the gross length of iron, 22 cms., is sufficient.
If the density in the teeth had come out too high, we should have had either to
lengthen the machine or take a different size of slot, and possibly a different number
of slots with a different number of conductors.

492 DYNAMO-ELECTRIC MACHINERY

In order to calculate the loss in the iron behind the slots we must find the working

flux per pole. This is

0256xl0«x0-685^^.3^^^^
4

The cross-section of the core is 165 sq. cms., giving a flux-density of 13,200.
This (from Fig. 29) gives a loss of 0 034 watt per cu. cm. As the total volume is
13,500, we have the iron loss behind the slot equal to 460 watts. This added to
380 gives 840 watts total iron loss. The loss on the part of the copper conductors
of the armature which are buried in the slots (as calculated below) amounts to
680 watts, so that the total watts to be dissipated by the iron surfaces of the armature
are 1520.

The number of watts which can be dissipated from the surface of an armature
ventilated in any particular manner can only be ascertained by trial, and the ratings
of all machines of this kind are really based on experiment. However, to illustrate
the methods of calculating the temperature rise given in Chapter X., we will apply
them to this machine. It will be found in practice that they give a very fair indica-
tion of what the temperature rise will be.

The most important cooling surface is the cylindrical surface of the armature.

The velocity is 17 metres per second. Allowing 35° C. for the difference between

the iron and air, we have

op, 333 X wts. sq. cm.

^^ ITTt '

Watts per sq. cm. =0-285.

As the cylindrical siirface is 3020 sq. cms., we can get rid of 0-285 x 3020 = 860
watts.

The velocity of air in the ducts of these little machines is always an uncertain
quantity, because it is so much affected by the obstruetions in the path of the air.
However, in a small armature with wide teeth and vents, not less than 0-75 cm.
wide, it may be taken at jth the peripheral speed. In this case, say 3 metres per sec.

Prom page 242, we have ^„=0-0014x 3=0-0042; allowing 20° difference of
temperature between iron and air in the vents, we have

20 X 0-0042 X 5000 = 420 watts dissipated by the sides of the vents.

The outside of the end plates and the inside cylindrical surface present a cooling
siirface of 2300 sq. cms. Allowing 0-15 watt per sq. cm. (see p. 254), we get rid of an
additional 350 watts, making the total 1630 watts for 45° C. rise. As the total watts
to be dissipated by these surfaces are 1520, we are well within the guarantee.

We now come to the compartment of the calculation sheet marked " conductors."
The throw of the coils will be 1 and 11, the pitch being 10. There are 12 conductors
per slot, making 492 in all, that is, 246 in series. The size we have already dealt
with. The length in the slots is 22 cms. and the length outside can be found from
the drawing, or, where no drawing is at hand, from the formula (1 -4 x pole pitch) + 5.
This gives us 54 cms., so the length of one conductor and its end connector is 0-76
metre. Multiplying by the number of conductors we get a total length of 374 metres.
Multiplying 890 by 0 17 (see page 143) we get the weight of a 1000 metres = 150
kilograms, so that 376 metres weigh 56-6 kilograms. The resistance of 1000 metres

CONTINUOUS-CURRENT GENERATORS 493

is found by dividing 0 17 by the section. This gives us just 1 ohm per 1000 metres,
so that the resistance of all the conductors in series is 0-374 ohm. As there are two
paths in parallel, we divide by 4 and get 0 094 ohm at 15° C. Now consider the
cooling conditions. One metre length of coil containing 12 conductors, when hot,

will cause 12 x 0 001 x 1 -2 x 72 x 72 = 75 watts loss.

The cooling surface of this metre length of coil will be 960 sq. cms., so the
watts per sq. cm. = 0 078. As the thickness of the insulating wall is 0 • 15 cm., we have

015x0078

00012

= 10°C.

difference of temperature between the copper and iron of the armature. This is
not too much. Where the watt* per sq. cm. of cooling surface of coil are below
0 08 watt per sq. cm., the cooling conditions of the end -windings, over which the
air is forced by the fan, are sufficiently good.

Ampere-tums per pole. We may conveniently work out the ampere-turns
per pole for three different voltages, 525, 540 and 560 volts.

In a machine of this size the length of the air-gap is fixed from two considerations.
In the first place, it must not be so small as to leave any danger of the armature
coming in contact with the field poles after the bearings are somewhat worn. In
the second place, it must be sufficiently great to prevent excessive distortion of the
field by the armature magnetomotive force. In a machine having no compensating
winding, it is desirable to have the ampere-turns on the field a little more than the
armature ampere-turns per pole. In this case the armature ampere-turns per
pole are equal to 4420. The field ampere-turns ought not to be less than 5200 at
full load. A rough preliminary calculation shows us that the ampere-turns on the
teeth and other parts of the magnetic circuit will amount to about 2500 ; and
allowing 20 per cent, of the armature ampere-turns for the increase between the
no-load and full-load excitation, say 900 ampere-turns, we should have on the
air-gap about 2100 ampere-turns. The density in the gap at 540 volts is obtained
by dividing 0-256 x 10® by the gap area 3020. This gives us the flux-density in the
gap of 8500. A rough preliminary calculation again gives us the length of gap at
about 0-254 cm. ; and taking this figure, we proceed to work out the magnetization
curve. First note that -K/is in this case the same as Ke, viz. 0-685.

TP, , 0-256x108x0-685 . ^^ ,^
Flux per pole = j = 4 -35 x 10*.

The leakage, worked out by the method given on page 326, is equal to
0-66 X 10« at no load and 0-8 x 10« at full load.

In working out the magnetization curve, we will take first the volts as 540.
The density in the core is obtained by dividing 4-35 x 10® by 2 x 165 sq. cms., and
is equal to 13,200. This requires 10 ampere-turns per cm., giving 100 ampere-turns
on the core. The density in the rotor teeth has already been worked out at 21,000,
and requires 350 ampere-turns per cm., or 1400 ampere-turns on the teeth. The
gap coefficient Kg\am this case 1 -25, so that with a flux-density of 8500 and a gap
length of 0-254 cm., we have the ampere-turns on the gap equal to

8500 X 0-254 x 1 -25 x 0-796 = 2150 ampere-turns.

494 DYNAMO-ELECTRIC MACHINERY

For the reasons given below, we will take a cylindrical pole body made of good
iron ; and as tbis may be worked at a density of aboutl6,000 C.G.s. lines per sq. cm.,
it may have a cross-section of about 314 sq. cm. The length of the pole body will
depend upon the dimensions of our standard frame ; in Fig. 427 we find it to be
13 cms. The ampere-turns per cm. at no load will be about 40, giving about 520
ampere-turns on the pole. The length of the yoke (see Fig. 427) is about 42 cms. ;
this is made of rolled steel ingot bent to shape, and has a cross-section of 192 sq. cms.,
giving a flux-density at no load of 13,400, requiring 9*5 ampere-turns per cm. This
gives us 400 for the yoke. Thus the total ampere-turns per pole at no load are
4570 at 540 volts. In order to plot the magnetization curve, it is generally sufficient
to take two other voltage points, say at 525 volts and 560 volts. The flux-density
in the various parts are then best found from the slide-rule, being approximately
proportional to the voltage. The form on page 489 gives the results. Thus we
have 3841 ampere-turns per pole at 525 volts, and 5397 at 560 volts. In plotting
the magnetization curve it will be found most convenient to take as ordinates the
flux per pole instead of the voltage ; so that in making calculations on the same
frame (the number of conductors in the armature being such as to give the required
voltage), we obtain the ampere-turns on the pole from the magnetic loading of
the frame direct.

In a continuous-ciirrent generator, even when the brushes are placed upon the
'* neutral," it is found that it is necessary to considerably increase the ampere-
turns at full load over the ampere-turns at no load in order to keep up the voltage.
This increase is due in the first place to the resistance of the armature and brushes
and of any series windings, and in the second place to the supersaturation of the
teeth under the trailing horn, brought about by the cross-magnetization of the
armature. Where a machine is fitted with commutating poles, the question whether
there is any demagnetizing effect of the armature depends upon the exact position
of the brushes on the commutator, and on the strength of the commutating pole.
Where the strength of the commutating pole is such that commutation takes place
behind the no-load neutral plane, the armature will have (on a generator) a magnetiz-
ing effect instead of a demagnetizing effect, and the extra ampere-turns put upon
the field magnet in this way may be made to compensate for the drop in voltage
which would otherwise be caused by the cross-magnetization effect and consequent
supersaturation of the teeth.

As it is always possible, after a machine comes on test, so to adjust the strength
of the commutating pole and the position of the brushes as to get a sufficiently
small drop in voltage between no load and full load, an exact calculation as to the
amount of extra ampere-turns required on the field pole to compensate for the
supersaturation of the teeth on load is not usually necessary. A generator with
its brushes rocked too far back will not run well in parallel with another generator,
so that it is not advisable to depend too much upon this compensating effect. It
is well to allow for an increase in the field-turns of some 10 per cent, on full load.
With an armature such as we have under consideration, the ampere-turns in
which are 0-9 of the field ampere-turns at no load, and in which the teeth absorb
nearly one-third of the total ampere-turns on the pole, the increase in the ampere-
turns at full load may be taken to be about 15 per cent, of the no-load ampere-turns.

CONTINUOUS-CURRENT GENERATORS 495

Thus, to obtain 540 volts generated (that is, 525 volts at terminals), we will require
5200 ampere-turns per pole. It is well to design the shunt coil so that it will take
continuously this full excitation, instead of relying upon the rocking back of the
brushes to supply the extra ampere-turns needed on full load.

Having decided upon the maximum number of ampere-turns required on the
shunt coil, the number of turns of wire will be settled from one of two considerations :

(1) We may wish to build the generator as cheaply as possible, using the smallest

amount of wire that will give us a temperature rise not* greater than the
guaranteed temperature rise.

(2) We may be ruled by considerations of efficiency and settle the number of

watts which are to be wasted in shunt excitation.

A large number of buyers will buy the cheapest machine that appears to be
good enough for their purpose. Other buyers, on the other hand, recognize that
very often a more expensive machine of higher efficiency will save more in the year
than the interest on the extra cost. With power at one halfpenny per unit, a
kilowatt for twelve hours a day for 300 days in the year will cost £7. lO^. per annum.
Capitalizing this at 10 per cent., we get £75. It would in many cases be worth
while for a buyer to pay £75 more for a machine which will save him 1 kilowatt
in the shunt excitation.

In the machine worked out on page 489, the loss in the shunt coils and rheostat
is 820 watts. The weight of copper is 64 kilograms. This is almost the minimum
weight we could use if we are to meet the temperature guarantees. It would be
good policy to increase this weight and make a saving in shunt losses if the buyer
would recognize the fact, and pay a greater price. There is room for another 1500
turns, which would reduce the losses by 250 watts. This, on the above basis of
calculation, could be capitalized at £19, and yet the cost of the extra 1500 turns
would not be more than £8. Yet so keen are many buyers to buy the cheaper
machine, heedless of small differences in efficiency, that the practice of using the
least possible quantity of copper pays from the manufacturer's point of view.

The same want of regard on the part of the buyer for small differences in efficiency
leads many manufacturers to use ordinary dynamo steel of good quality at (say)
£11 per ton rather than alloyed steel at £25 per ton. In the present case, with
ordinary iron the iron losses work out at 840 watts, whereas with alloyed steel they
could be reduced certainly to 600 watts. The saving of 240 watts is worth about
£18, and the cost of the alloyed iron would not be more than £5. Some buyers are
beginning to recognize these facts, and the future may see a very great increase in
the efficiency of small generators and motors.

Rectangular coils versus circular coils. A good deal of discussion has taken place
between designers on the merits and demerits of coils wound on a cylindrical former
and coils wound on a rectangular former. The advantages of the cylindrical coil
are as follows :

(1) The length of turn for a given area enclosed is only 0-89 of the length of a

turn for a square coil, and a smaller fraction still of a turn of a coil whose
length is greater than its breadth.

(2) It can be wound by means of a machine, so that the labour in winding is

considerably reduced.

>6 DYNAMO-ELECTRIC MACHIKERY

(3) No insulation is required other than the cotton covering between layers ;

whereas with rectangular coils it is usual, when winding the wire " layer
for layer," to insert insulation at tiie cornere, in order to enable the wires
to be drifted over as each layer is put on.

(4) The cylindrical coil, when wound "layer for layer," can be made much

tighter aad more compact than is possible with a rectangular coil, and the
heat conductivity is therefore much increased.

(5) The bobbins are exceedingly clieap and easy to manufacture.

(6) The number of moulds to be kept in stock is reduced.
The disadvantages of the cylindrical coil are :

(1) It takes up more room measured along the periphery of the armature than

a rectangular coil enclosing the same area. It, therefore, does not allow
so much room between poles. This does not matter so much ou four-pole
machines on account of the great angle between the centre lines of the
poles. If the axial length of the machine is not more than 0-8 of the pole
pitch, the round pole limb leaves plenty of room for the insertion of a
commutating pole and winding.

(2) It is not so easy to change the axial length of a frame when it is fitted with

round poles. It is, however, possible to design a standard line of machines
with three or more economical axial lengths.
We have adopted the round pole and cylindrical coil for four-pole machines,
icause we believe that there is nothing to be gained by making the axial length of

Fia. 430. — Showing uringeineDt at rooDd aUel pale body and rectanEuUc pole shoe.

the armature greater than 0'8 of the pole pitch, and up to this length the round pole
can be used. In laying out a standard line of frames, there might be three dtSerent
a.Yial lengths for armatures 43-4 cms. in diameter: 27 cms., 22 cms. and 17-5 cms.
For these, three diSerent diameters of round poles could be used, 22 cms., 20 cms.
and 18 cms. The same punching for the pole shoe can be used in all cases, built
up to different lengths. The punched pole shoe (see Figs. 429 and 430) is secured
to the pole as follows : After building up the shoe and riveting together by means of

CONTINUOUS-CURRENT GENERATORS

497

two axial rivets, four suitable points are chosen on the face which is to lie adjacent
to the pole limb, and the iron of the punchings at these four points is melted together
by means of an oxy-acetylene flame. These four points are then drilled and counter-
sunk to receive screws which are screwed into the pole limb. On the pole limb
22 cms. in diameter it is necessary to use a built-up winding of partly conical form ;
but the other two take plain cylindrical coils which are exceedingly cheap to
manufacture.

Yoke. The yoke may either be of cast steel and be cylindrical in form, as shown
in Fig. 431, or it may be made of rolled ingot bent into a cylindrical or octagonal

Fio. 481. — Showing arrangement of circular yoke of cast steel for 75 K.w. CO. generator.

shape.. Where a large number of yokes are made in a forge provided with suitable
machinery, the labour of bending small yokes into shape is not excessive. A steel
casting for the 75 K.w. generator under consideration would weigh 8| cwt., and at
138. per cwt. will cost £5 128, before any machining is done on it. The metal of
the yoke shown in Fig. 429 rolled roughly to size will weigh 71 cwt., and at 9«. per
cwt. will cost £3 7s. There is more than enough difference in the first cost of
material to pay for the forging pf the frame. In the octagonal frame (Fig. 429) the
machining of the surfaces to receive the poles can be carried out with a pin-cutter
mounted on the tool that drills the sockets for the poles. The only other machining
is on the faces where the two halves of the yoke meet, and the turning of the ends
to receive the cast-iron end brackets. It will be seen that the feet of the generator
are cast with the end brackets.

W.M. 2 1

498 DYNAMO-ELECTRIC MACHINERY

Shnnt winding. As pointed out above, the shuDt coil has been worked out for
the minimum weight of copper. Beginning with the ampere-turns required at
full load, we make a preliminary estimate of the total cooling siirface of the coils.
This is 10,000 sq. cms. Allowing 14 sq. cms. per watt, we are able to get rid of
720 watts. The approximate voltage expended on the shunt coils is found by
deducting about 50 volts from 525 to allow for some margin on the rheostat. Divid-
ing this voltage into the 720 watts, we find the approximate exciting current and
from it the number of turns. The mean length of turn can then be found approxi-
mately, and the total length of wire, and hence the resistance cold and hot. Having
chosen a wire which gives us approximately the right resistance, we can go over
the figures again and get the values as shown in the calculation sheet.

Gommatatmg pole. In order to calculate the flux-density B<. under the com«
mutating pole required to bring about commutation, we proceed as indicated on

page 480. , / 3 o-5\

Z„=1.25(3-^4:2 + ?|) = l-57.

i:.=0-46x^fg(log,o i|-0-2) = l-32,

If the axial length of the pole tip is 14 effective cms., and the length of the gap
under the commutating pole is 0*41 cm., we require

22
2265 X Y2 X 1 '25 X 0-41 X 0-796 = 1450 ampere-turns per pole.

Add to this the armature ampere-turns of 4420, and we get about 5900 ampere-
turns total. We will therefore require 41 turns, carrying 144 amperes.

The width of the pole is 3 cms. at the tip, and as the air-gap is 0-41, we will have
a fringing field, which, in conjunction with the short-throw coil, will give a diminishing
commutating E.M.F. towards the end of the period of commutation (see Fig. 423).
For this reason we have taken the coefficient 2 -8 instead of 2 in the formula given
above. It will be seen from the drawing (Fig. 429) that we have made the com-
mutating pole 6 cms. wide at the root so as to avoid saturation on considerable
over loads. By keeping the base of the pole wide we are able to shorten the axial
length, and thus to save a great amount of copper in the coil. It is in fact quite
good practice to make the commutating pole of round section in cases where sufficient
room can be found for the rather wider limb required in this case. If a 60 K.w.
generator be built upon the same frame, but with 17*5 cms. length of iron instead
of 22 cms., it will be found that the diameter of the main pole will be reduced, and
this gives room for a round commutating pole of ample section.

The commutator. In designing a commutator for a small machine of this kind
simplicity and economy are important. There is no great danger from expansion
troubles such as occur on large commutators, so that it is sufficient to support the bars
between V-rings, one of which is turned on a cast-iron bush which forms the main

CONTINUOUS-CURRENT GENERATORS 499

support of the commutator, the other being a drop-forging pressed in by means of
a screwed washer. The bars will, of course, be of drawn copper and the insulation
of mica.

In the design under consideration, we have 123 bars, or 31 bars per pole. This
is as great a number as one can economically provide on a small machine. . The
number is found to be amply sufficient where the current to be collected is only
144 amperes.

Width of brashes. In settling the width of brush to be used on a commutating
pole machine, one must have regard to the length of arc over which the short-
circuited coil travels before the short circuit is removed. As long as this arc lies
well away from the horns of the main pole, the brush is not too wide. One m&j
allow it to extend within such a distance of the on-coming pole as to have it moving
in a field from that pole almost equal to the field of the commutating pole. We may
then have quite good commutation at full load, but as the field of the main pole
gets weaker on load and stronger on no load, it is advisable to shorten the arc so
that it lies almost entirely under the influence of the commutating pole. If in the
last moments of the commutating period the field strength is reduced (see Fig. 423),
the adjustment of the commutating winding will be found somewhat easier. Within
these limits the wider the brush used on a conmiutating pole machine the better,
as the time of commutation is increased and the e.m.f. required is smaller. In
the machine under consideration we have made the brush 2 cms. wide. This
makes ^«-l-&p-Cp = 4'65 cms., that is to say, just 1*65 cms. wider than the tip of
the commutating pole.

As we have put two brushes per arm with an area of 18 sq. cms., the amperes
per sq. cm. are only 4. This is lower than it need be. A density of 6 would do,
but it is not always possible to fit in standard brushes so as to give the most econo-
mical arrangement.

The cooling surface of the commutator works out at 1000 sq. cms., so with 300
watts lost we have 0-3 watt per sq. cm. There is no danger of overheating if we
are not troubled with high mica or some other cause of bad contact between com-
mutator and brushes. It is good practice to mill out the mica to a depth of 1 mm.
Brushes of ordinary hard carbon are reconmiended on this machine.

Efficiency. The way of working out the efficiency is sufficiently clear from the
calculation form. The windage is considerably increased by the addition of the
fan shown in Fig. 429. A machine of this kind will have a friction and windage
loss of about 600 watts without the fan, and about 1000 watts with the fan.

It is very desirable to see that the fan is not very much greater than is necessary
for the purpose of keeping down the temperature. If the machine runs much below
the guaranteed temperature rise, it should be rated for a higher output or the fan
should be reduced so as to lower the windage losses.

The figures for the iron loss have been increased a little on load to aUow for the
increased losses on the teeth. The field losses taken should include the losses in
the rheostat. The PR losses include those in the armature and in the conmiutating
pole winding. The brush losses are taken as if the voltage drop in positive and
negative brushes amounted to 2 1 volts. This is justified by the low current-density.
The total losses at full load are 5-77 K.w., giving an efficiency of 92-8 per cent.

600

DYNAMO-ELECTRIC MACHINERY

Characteristics
of Oenerator.

SPECIFICATION No. 11.

1000 K.W. CONTINUOUS-CURRENT GENERATOR TO FORM PART

OF A MOTOR-GENERATOR SET.

155. This specification provides for the supply, erection,
testing and setting to work of a continuous-current generator
having the following characteristics :

Normal output
Voltage adjustable be-
tween
FuU load current
Speed
How driven

Temperature rise after
6 hours fall-load run

Over load

Temperature rise after
30 minutes over load

Puncture test

{

1000 K.w.

460 and 500.

2000 amperes.

246 R.P.M.

Direct connected to induction

motor.
45° C. by thermometer.
50° C. by resistance.
2300 amperes for 30 minutes.

55° C. by thermometer.

1500 volts (alternating) applied

for 1 minute between wind-

ings and frame.

Excitetion. 156. The generator is to be shunt wound.

Duty.

157. The generator is intended to supply continuous current
for general Ughting and power work for the Town of
It is intended to run in parallel with other continuous-current
shunt-wound machines, some of which are motor-driven and
some steam-driven. The particulars of these machines are
given in Schedule I.

158. The contract will include the deUvery of the generator,
together with bedplate, half-coupUng, bearing, and pedestal,
at the sub-station at ; and the erecting,
aUgning and coupUng of the same to the 1500-h.p. motor
described in Specification No. . The switchgear and cable
work are provided for under another specification.

Foundations. (See Clauses 6, p. 271 ; 36, p. 360 ; 74, p. 382 ; 272, p. 591.)

Extent of
Work.

CONTINUOUS-CURRENT GENERATORS 601

159. The frame shall be split horizontally and arranged Horiiontaiiy
so that the armature may easily be inspected and lifted out '^
without dismantling the brush-gear.

160. The generator shall be of the ordinary multipolar type Type of
with drum-wound armature. The armature coils shall be
placed in open slots and held so that they can be readily
renewed.

161. The commutator shall be of ample proportions, con- commutator.
structed according to the best practice. It shall be thoroughly
seasoned before delivery, and after having been ground true

once on site shall not show any signs of high bars, high mica
or appreciable eccentricity. The mica may be cut out for
^ inch below the commutator surface if the Contractor will
guarantee that no dirt will lodge in the grooves so made,
under the conditions of running experienced in the sub-
station in question. The wearing depth of the commutator
shall not be less than f inch.

162. The brushes shall be of ordinary carbon, and thesniahes.
commutating conditions shall be such that good commutation

can be effected without resorting to some special type of
brush. A sample brush with its market price afiixed shall
be supphed with the tender.

163. A sample brush-holder shall be supplied with the Bruah-hoider.
tender.

164. The generator shall run sparklessly at all loads up commutauon.
to 26 per cent, over load at any pressure between 460 and

600 volts.

166. The Contractor shall state the drop in voltage between Eeguution.*
no load at 250 r.p.m. and full load at 246, which he proposes
to give in order to run in parallel with the generators set
out in Schedule I. He shall also state the rise in voltage
which will occur when load is thrown off. This change in
voltage shall not be. more than is necessary for parallel
operation, as it is desired to obtain the best possible regulation
on the sub-station.

* Where the generators in the sub-station are compound-wound, particulars
should be given of the actual rise in voltaffe between no load and fuU load on the
sub-station. It should be stated whether the series coils are to be connected on the
poeitive or negative side of the generator, and particulars should be given of the actual
voltace between the equaliser bar and the positive or negative bar, as the case may be,
with nill load on the sub-station : that is to say, of the voltage upon the series windings
of the generators at present installed.

502

DYNAMO-ELECTRIC MACHINERY

£fflciency.

Bheostat.

Tests before
Shipment.

Of Besistanoes.

Magnetization
Curve.

Short Circuit.

166. The efficiency shall be detennined from measurements
of the separate losses. The iron loss at 500 volts, the friction
with all Drushes adjusted for their workingpressure, and the
windage, shall be measured at no load. The resistances of
the armature and commutating winding shaU be measured
at a known temperature ; and the PR loss calculated at
60° C. The drop in the brushes shall be taken to be 2-3
volts for the purpose of calculating the brush losses. The
field and rheostat losses shall be taken as together equal to
the product of the amperes of field current at full load at
500 volts into the voltage. All the above losses, expressed
in kilowatts, shall be added to the kilowatt output, and the
ratio of the output to this sum shall be taken as the calculated
efficiency. The Contractor shall state in the Schedule*
attached the efficiency of his generator calculated in this
way at full, three-quarter and half load at 500 volts, and he
shall guarantee that there shall be nothing in the construction
of the machine that will lower the actual efficiency when
running on load by more than 15 per cent, below the figures
so given.

167. A field rheostat with multi-contact switch is to be
provided in the field circuit of the generator, of sufficient
capacity to lower the voltage of the armature to 460 volts
at no load when the machine is cold, and to enable the voltage
to be raised to 500 volts when the generator is deUvering
1250 K.w. in the hottest weather. Sufficient contacts must
be provided on the switch to make the voltage change very
gradually as the switch is moved over the whole range.
One step of the rheostat must not change the voltage by
more than 1-5 volts at any load and at any part of the range
when the machine is operating by itself.

168. The following tests shall be carried out at the maker's
works before shipment :

(a) Measurements shall be made of the resistance of
the armature and field windings.

(&) The generator shall be run at full speed, no load,*
with the field excited, and measurements shall be taken

* In some cases, where it is impossible to carry out a full-load test, the Purchaser
may require to have a full-current commutation test on short circuit. The clause
calung for this can be worded as follows :

(6'). The generator shall be run at full speed with the armature short-circuited
through the commutating pol^s and an ampere-meter, the field windings being excited
BO as to give full-load current ; the machine shall, under these conditions, commutate
well. Measurements shall be taken of the temperature rise of the commutator and
armature after six hours* run.

CONTINUOUS-CURRENT GENERATORS 503

showing the relations between field current and voltatge
generated, the iron loss at various voltages, and the iron loss.
friction and windage.

(c) The generator shall be run at full field-current Fiew Heating
for six hours, and measurements taken of the field resist-
ance while hot.

(d) While the machine is still hot, an alternating Puncture rest,
pressure of 1500 volts (virtual) shall be applied between

the armature winding and frame for one minute.

The following tests shall be carried out after erection on Tests after

ii 'j i» • -t Erection.

the site aforesaid :

(e) After erection on site, the generator shall be run Temperature
at fall, load for six hours, and for two hours on the stated
overload ; and measurements shall be taken of the
temperature of the armature, the windings and iron, and

the field windings, by thermometer, and of the field wind-
ings by resistance, to see that the specified temperature
rises above the surrounding air are not exceeded. For
the purpose of these tests, the temperature of the room
shall be taken three feet away from the generator in line
with the shaft.

(/) A test shall be made to ascertain the drop in Regulation.
voltage between no load and full load, the speed at no load
being approximately 250 r.p.m., and the speed at full
load being 246 r.p.m. Tests shall also be taken to ascertain
the rise in voltage between full load at 246 r.p.m. and no
load at 250 r.p.m., in order to ascertain whether the
guarantees given by the Contractor have been met.

{g) The generator shall be run on its ordinary daily Endurance
load for one week under the direction of the Contractor's
engineer, to see that all matters are in order. It need not
be accepted by the purchaser until it is complete in every
particular.

169. The Tenderer shall quote separate prices for the spares,
following spare parts :

(1) A field coil.

(2) Twelve armature coils.

(3) One set of brushes.

(4) Enough brush-holders to complete one brush arm .

504

DYNAMO-ELECTRIC MACHINERY

Date ^*/>^<rr,,/3.. TT^fifG CC CBW .r..S¥H ■'l^*?* ««»«^ W^'^- ^^ .„ Etoc. Sp«. .../.^

K^ mo . P F : >ha.e ; Volt, tf ^.<?.-4.<><> .; A»p. per ter.^^^;...; Cyde...^.^. ..; R-PM^^f ••; ^/S^ '^L"'

Customer : Order No.

Qnot No..

Perf. Spec ; Fly-wheel effect

^-^,C..cun.^75;G,pAre,/7<m^^xi^j>:

kZa

^po«.uz« ;Ja£* __^ ihukrpm ^^^5

W^l9200a ; Q\,com,.S3^ ; IC.V.A. ...fT

K. li.

Arm. A.T.p. pole...i3C<?.<?...

Apmatupe

o
o

I-

Dia. Outs

Dia. Ins

Gross Length
Air Vents — 3:

Opening Min

Air Velocity

Net Length-^
Depth b. Slots-
Section ^SO.-

Flux Density—

Loss:i^p. csy.CM. Total
Buried Cu.i--3^^-Total
Gap kxft^nopQ^ wts

Vent Area_fl.4PPi? . Wts
Outs. Area /^^<?^; Wts

No of Scgs /^JMn.Circ.
No of Slois I i^t\ X •P^=

K, J'Q^

Section Teeth .
Volume TeetlL.
Flux Density —
Loss;^ pcuiZ^Total

Weight of Iron —

3
"D
C
O

Star or Mesh Throw

Cond. p Slot ,

Total Conds 96jnserm

Size of Cond. -JZ x/.2-

Amp. p. sq._C^— ^— j

Length in Slots ->^^

Length outside ^^ Sura

Total I^gth ^-^^5

Wt. of ^ffio-^^-^- Total

Res. p. i.®io'4i^-Total

Watts p./ZU-^^ .

Surface p. OL^^M-S^n.
Watts p. Sq

'00/2

..Max. Fid. hT.SS.OQ.

Field 8Ut

Boxe
\ Total Air Gap
Gap Co-eff. K.

/84-

/•/

Pole Pitch 4? Pole Arc Lj^

Flux per Pole-/<>g^^<^*

Leakage n.I iXj^Z.

kxf^f90 Fhix density
Unbalanced PuH

•7/

/2'5x

I5900

No.ofSeg.
No.of Slots

Vents

K. -^

MzlCiic.

.Section

Weight of Iron 'O pof^^

'TSSo

Borloo Oomm.

A.T. p Pole n.Load _i!2!f:2

A.T. p. Polef .Load^.55£-l.

i2J0OO

Surface
Surface p. Watt.

I*. R

LR.

Amps. .

No. of Turns —
Mean 1. Turn —
Total Length
Resistance

gfiggp.U

MZQQ
iO'S

360,

SOOO

2000

600

raz

i 9SOO\

i9'r/coMy24,'m^

Res. per i.ooo l£iS&

j3S.

0Qff7S

Size of Cond.

0/7

an. \t0s9Cms.

Conds. per Slot.

Total

Length

Wt per i.oool

Total Wt

Watts per Sq..
Star or Mesh .

6900

70O ^i/ogrs. 3SOJ^

*062

Paths in paralld

.i//2. Volts.

A-Lp*-* a T.

2LQ0S_39Q1 2000^220^
4720
900

fS90O\_3p

i :7ac?4|

.S4:0.yo\XA,

/630C
93<K

A.T.l>-tm,
r -J —

2i£Q,

A.T.

5000
/SOO

2/0

9/60

Commutator.

Volts p. Bar /f '7 —
Bis. p. Arm

Size of Bxs. ^^^^

Amps p. sq.C^. ^'?-~,^
BvSiU^46Q0+$0Q0

Watts p. Sq

Mag. Cur. Loss Cur.

Pcnii . i8tftt,Stot £ne/s / ' 74

Output -
Input —
Efficiency

/5/fl i I OS 3^
'\94-SC95_

"iTT I —

2 X

177

End

79/ \J32_
94^-6 \ 93-9!

Rot. Slot X
Zig-zag

X
X X

Amps ; Tot

; X. =

/'73

'94.

.44/

« +

Imp. V +
Sh. cir. Cur.

Starting Torque
Max. Torque -
Max. H.P

SUp

Power Factor

CONTINUOUS-CURRENT GENERATORS 505

THE DESIGN OF A 1000-K.W. C.C. GENERATOR TO MEET

SPECinCATION NO. 11.

460-500 volts ; 2000 amperes ; 246 R.p.m.

The procediire in designing this machine follows very closely that adopted for
the small machine given on page 488. The L^l constant for large multipolar c.c.
generators fitted with commutating poles is, however, much smaller than for little
four-pole machines. It will be found that a D^l constant of 2*4x10^ will give
us a frame not too small to meet the temperature guarantee. The calculation
sheet is given on page 504.

The dioice of the number of poles is a matter of importance. The considerations
which settle the number of poles are as follows : Where the current to be delivered
is very great, the number of brush-arms will be increased until the current per
brush-arm is not excessive. Thus, in generators for electrolytic work, delivering
many thousands of amperes at a low voltage, one may take 1000 amperes per
brush-arm as a suitable figure, and fix the number of poles accordingly. When the
voltage is higher, say 250 volts, a rather lower current per brush-arm will generally
be chosen, from 500 to 750. On 500-volt machines it is usual to choose a still
lower current per brush-arm, say from 300 amperes for machines of 500 K.w.
capacity, up to 500 amperes for very large generators. It is often worth while to
work out two or three designs with varying numbers of poles to see which arrange-
ment makes the cheapest good machine on the available frames. In this case we
have to. deliver 2000 amperes, so that a twelve-pole machine would have 333
amperes per brush-arm. No advantage is to be gained by reducing the number of
poles, as this would only increase the length of the commutator and the axial length
of the iron. In this respect a c.c. generator whose speed is prescribed differs from
a rotary converter, in which the diminution in the number of poles increases the
speed and brings about a saving in the material. In actual practice, the diameter
would be fixed by the diameter of some frame which the manufacturer might have
developed ; but if we were starting de novo we should have to make a compromise
between building a machine of large diameter and short axial length, which would
give us good commutating conditions, and building a machine of smaller diameter
and greater axial length, which, though economical in material, might give us an
excessive inductance in the armature coils. A happy mean is generally to be found
in making the pole of the generator approximately square in section, or, as is some-
times preferred, somewhat longer in an axial direction than in a circumferential
direction. In this case, if we take a square pole 29 cms. x 29 cms., we find that
the diameter is just great enough to enable us to get in the requisite number of
conductors.

The number of conductors is controlled by the number of commutator bars
which we wish to have per pole. From the considerations given on page 532, we
will decide on 48 bars per pole ; so that in a lap winding we have 96 conductors in
series. On these large multipolar machines fitted with commutating poles there is
no difficulty in obtaining a coefficient Ke (see page 13) as high as 0*71. Adopting

DYNAMO-ELECTRIC MACHINERY

FlOB. tSe and 433.— Sactloaal TlewB ol 1000 K.kt. c.C. gf nvrskir, &00 volts, 24fl R.

CONTINUOUS-CURRENT GENERATORS

507

and 1 : 4. This generator and the motor illustrated in Fig. 407 form a motor-generator set.

508 DYNAMO-ELECTRIC MACHINERY

this coefficient, and allowing 10 volts for drop of voltage in the armature, we find
the value of Ag^ from the equation

510=0-71 X 41 X 96 x-4^B.
Hence AgS = 1 -83 x lO®.

If we take a diameter of 183 cms. and an axial length of 29-6 cms., we have a
circumference of 575 cms. and an area of gap of 17,000 sq. cms. We make a check
calculation at this point by dividing AgB by the gap area to see that B is somewhere
in the vicinity of 10,000 c.G.s. lines. In this case B in the gap will equal 10,750,
which is not too high a value for a c.o. machine if we can give enough area to the
section of the teeth. We must also make a check calculation to ascertain the ampere-
wires per cm. of periphery. The total ampere-wires laZ/a will equal 192,000, and

laZa

circumference

=334.

This is a suitable figure for a large machine, and will permit us to work the
copper at approximately 450 amperes per sq. cm.

The amperes per conductor are 166. The cross-section of the conductor may be
taken as 0-375 sq. cm., and we may fix on a copper strap 0-2x1-9 cms. This, with
insulation and suitable room for the retaining wedge (see Fig. 158), will require a
slot 0-96 cm. wide x 5-1 cms. deep.

The machine is illustrated in Figs. 432 and 433.

The next step is to check the saturation in the teeth.

^(183 -7) =555.

This gives us the mean circumference of the circle through the teeth. Subtracting
from this 184, the width of all the slots, we get 371, the width of all the teeth. The
net length will be 23 cms., giving us a cross-section of all the teeth of 8540. The
apparent flux-density will therefore be 1-83x108-^8540=21,400. From Fig. 46
we see that the actual flux-density will be 21,000. As the frequency is only
25 cycles, the loss (see page 52) will be 0-08 watt per cu. cm., giving 3-5 k.w.
loss in the teeth. The loss behind the slots and the buried copper loss are calculated
in the same manner as described on page 323. We find that the total watts dissipated
by the iron surfaces of the armature are 14,800. With a 45** C. rise we see, from the
calculation given in the sheet, that we can dissipate 16,800 watts. The calculation
of the watts per sq. cm. on the surface of the armature coils gives us 12^ C. difference
of temperature between copper and iron.

Magnetization curve. It is convenient to work out the ampere-turns per pole
at 460, 510 and 540 volts, as shown in the calculation sheet. It will be seen that,
owing to the high saturation of the teeth, the ampere-turns dn the teeth at 510
volts are 2000. The length of the air-gap will be adjusted so as to make the shunt
ampere-turns on the pole somewhere about equal to the armature ampere-turns
per pole. In this case, an air-gap of 0-5 cm. gives us 7804 ampere-turns per pole,
which is sufficiently near 8000 to prevent undue field distortion. At full load we
must allow for some further increase in the shunt ampere-turns ; on a commutating-
pole machine it is sufficient to add about 15 per cent, of the armature ampere-turns,
which will give us 9000 ampere-turns per pole to be provided at full load if the

CONTINUOUS-CURRENT GENERATORS

509

brashes are rocked slightly ahead of the neutral. In calculating the shunt winding,
we first make an estimate of the total cooling surface (see page 331), which is usually
taken from previous machines built on the same frame, or it can be found by trial
and error. In this case we have 86,000 sq. cms. ; allowing 16 sq. cms. per watt, we
have a permissible loss of 5400 watts. As it is desirable to have some margin in
our rheostat, we will take about 360 volts drop in the winding, so that the shunt

fnches 7 6

and Kapp Lines per sq. in.

lifito ^iooo tiaoo ahoo e^bc^ 3fioo J

C O S Lines per sq. in

FIG. 434. — Construction for calculating leakage flux between poles of 1000 K.W. c.c. generator.

amperes will be about 15. Dividing this into 9000, we get 600 as the approximate
number of turns. The main length of turn is 1-23 m., giving a total length of
9500 m. The approximate hot resistance can be obtained by dividing 360 by 15.
From this and the length of wire we find the resistance per 1000 metres, and then
the size of conductor, which must finally be adjusted to fit some standard size. It
is then a simple matter to run over the figures and adjust them more exactly.

The leakage between poles is calculated in the same manner as described on page
326. The graphic construction is given in Fig. 434.

510 DYNAMO-ELECTRIC MACHINERY

Commatating pole. It will be seen from the calculation sheet that the armature
coil lies in slots 1 and 12 ; so that it is short chorded by one slot. The effect of
this will be to slightly reduce the self-induction ; but the main advantage lies in
its giving a commutating curve of the type shown in Pig. 423, page 477. For the
purpose of getting a commutation curve of this type, the width of the commutating
pole must be made about the same as the pitch of the slot — ^in this case 3 cms. The
following is the calciilation of the various leakage coefficients :

X. = 1 -357 r^)= 0-94,

B<, = 2.8x4.41x3-|;^lf 5-3^ = 2120.

If now we make the effective axial length of the commutating pole, 14 cms.,
instead of the full axial length of the armature, 29-6 cms., B,. must be increased
to 4440. If the air-gap imder the pole be made 1 cm. long, the effective ampere-
turns upon the pole must be

4440 X 1 X 1 1 X 0-796 = 3900.

Six turns per pole multiplied by 2000 amperes would give us 12,000 ampere-turns
= 8000 + 4000.

A suitable diameter of the commutator would be 107 cms. ; this gives us
a circumference of 335 cms. with a pitch of brushes of 28 cms. A good way of
arranging the commutator V-rings and holding bolts is shown in Fig. 515. With 576
bars we have an average of 14*7 volts per bar ; with 6 brushes per arm, each measur-
ing 2x4-5 cms., we get a current density of 6-2 amperes per sq. cm. If we allow
2-3 volts drop on the positive and negative brushes, we have a resistance loss of
4600 watts ; and with a brush pressure of 2J lbs. per brush we have a friction loss
of 3000 watts : making a total loss in the commutator of 7600 watts. It will be
seen from Fig. 433, page 507, that the commutator is provided with very long lugs,
so that no difficulty will be experienced in dissipating the loss. The method of
working out the efficiency will be clearly seen on the calculation sheet.

SPECIAL CC. GENERATORS.

Before passing on to consider c.c. turbo-generators, we will take up a few matters
which arise in connection with very slow-speed machines and those which for some
reason cannot with advantage be built with ordinary lap windings.

We have seen that the fewer the number of turns per coil between two successive
bars on the commutator, the easier are the commutating conditions. On all large
machines we aim at getting only one turn per commutator bar. On small machines
of ordinary voltage we are compelled to have more turns per bar, because the
magnetic flux per pole is so small that we could not generate the voltage required

CONTINUOUS-CURRENT GENERATORS 611

without having several (and in very small machines many) turns per bar. The
number of bars between positive and negative brushes is limited by the fact that it
is not desirable to make the bars too narrow. For instance, on a commutator
9 ins. in diameter, we would not care to have more than 200 bars, or 50 bars per
pole on a four-pole generator. If now we must generate 500 volts, we have an
average of 10 volts per bar, or say 15 volts maximum, and on a small generator of
ordinary speed we would require several turns to generate the 15 volts.

For very small machines there is no great disadvantage in having a number of
turns per coil, because the current to be commutated is small ; but as we proceed
to 500-volt machines of 200 to 250 K.w. capacity, the commutation with two-turn
coils is not as good as we could wish, so it is better to resort to wave windings. A
two-circuit wave winding on a four-pole machine gives as many conductors in series
(for a given number of commutator bars) as a lap winding with two turns per coil ;
and although the voltage per bar is the same as for the lap winding, it has the
advantage of bringing about commutation of each single-turn coil separately, and
thus taking full advantage of the resistance of the carbon brush.

There are many cases, however, in which we caxmot with advantage make use
of the twocircuit winding. On machines of large size with many poles the voltage
per bar, with a two-circuit winding, becomes too great, and yet it may be that
a single-turn coil would not give us sufficient voltage. In these cases the Arnold
singly re-entrant multiplex winding is most useful. The need of a winding of this
kind is most commonly found in slow-speed continuous-current generators of
moderate output.

In order that we may fully appreciate the use of this winding, let us take a
500 H.P. 500-volt rolling-mill motor running at 32 R.P.M. Experience leads us
to a L^l constant of 3 x 10^ as suitable for a machine of this size. The armature
might have a diameter of 274 cms. and an axial length of about 50 cms. As it is
not economical to make the pole pitch too great, we might choose 18 poles,* giving
a pole pitch of 48 cms.

Now let us see what type of winding is best for such an armature.

We may have a flux-density in the air-gap of 9500, so that the AgB may be as

^8^ ^ IT X 274x 50x 9500=41 x 10«.

Taking K/ at 0-68 and allowing 30 volts drop in windings and brushes, we have

470=0-68 X 0-533 xZ,x 4-1.

Z«= about 320 conductors in series. Let us try a lap winding with as many cir-
cmts as there are poles. With only one turn per coil we would have 160 bars per
pole, which is clearly too many. With two turns per coil we would have 80 bars
per pole, still a large number. Three turns per coil would give us 54 bars per
pole, a suitable number ; but three turns per coil would not give us ideal commu-
tating conditions. ' '

* In this case the number of poles is not fixed by the amperes per brush arm, but rather by
the circumstance that the machme is very larae on account of its slow speed, and in a large
diameter many poles call for less material than fewer poles. At the same time, the small current
per brush arm does make the commutating conditions better than they otherwise would be, and
the small number of brushes per arm enables a narrow commutator to be used.

512 DYNAMO-ELECTRIC MACHINERY

Let ufl try the ordinary two-circuit winding. This would give us only 320 bars
on the whole commutator, or only 17-8 bars per pole. Moreover, the current per
conductor would be 415 amperes. The two-circuit winding is then out of the
question. Now try an Arnold multiplex singly re-entrant winding.

We will employ the following symbols :

2p = Number of poles.

2a » Number of armature circuits in parallel.
jffm= Number of commutator bars.
y = Throw on the commutator — ^that is, the number of bars between
one positive brush and the next positive brush.
i^r,=: Number of slots in the armature.

Then the quantities must ful^ the following conditions :

p must be a simple multiple of a.
Ng must be a simple multiple of a.
Kfn=-{y^p)±a.

y and Km must be prime to one another if the winding is to be singly re-
entrant.

Further, if we are given the number of blank stampings forming a circle in the
armature, Ns must be a simple multiple of the number of blanks.

Let us see how we can fulfil these conditions in the machine in question.

2jo=18,

/( the terminal amperes =830.

In choosing the number of parallel circuits we aim at making the current per
conductor somewhere between 125 and 250 amperes per conductor. If we divide
830 by 4 we would get 207 amperes, a suitable number for the amperes per con-
ductor, but 18 is not divisible by 4. We therefore try a=3, 2a = 6.

828

-w- = 138 amperes per conductor.

This is quite suitable. 2a x 3 » 18 = 2p.

Next, we have to settle on the number of conductors. This we can do by adopt-
ing an economical number of ampere-wires per cm. of periphery. This should be
between 250 and 360. Assume 300.

The circumference of the armature is 860 cms., and the current per conductor
138 amperes. Therefore a suitable number of conductors would be about

860 X 300 - o^n
-^38- = ^®^^-

And Km is half this number, or about 935.
Now apply the formula

Km=y^p±(i- j9 = 9 and a=3.

Try y = 107, a prime number.

Km = (107 X 9) ± 3 = 966 or 960.

960 is a more promising number than 966, because it will give us an even number
of slots. With 960 commutator bars we could have 240 slots with four bars per

CONTINUOUS-CURRENT GENERATORS 513

slot ; moreover, 240 slots is a likely number for fitting a possible number of blank
stampings per circle.

Now it will be seen that with 960 bars we satisfy all the conditions set out on
page 512. K„, = (107 X 9) - 3 = 960,

-y^=53-3 bars per pole,

ax3 = 9=j9,
ax80=iV, = 240.

107 and 960 are prime to one another, so we will have a singly re-entrant
winding with six circuits in parallel.*

The total number of conductors on the armature is 1920, so there are 320 con-
ductors in series. The last-given method of finding a suitable number of conductors
will not necessarily give us the same number as we found on page 51 1 by considering the
total flux of the frame and the speed ; but it will give us a number somewhere near it,
because the D^l constant is based on our working the frame at an AgB somewhere
about 4 1 X 10®, and the amperes per centimetre of periphery somewhere about 300.

Even in cases where we could make a passably good machine with a lap winding,
it will often be better to use Arnold's winding for the purpose of reducing the total
number of conductors on the armature. By reducing the number of paths in
parallel and increasing the current per conductor, and hence the size of the con-
ductor, we can save insulation space and make an arrangement in which the com-
mutating conditions are very good.

Take the case of a 200-K.w. 250- volt generator which has to run at the low speed
of 180 R.P.M. With a DH constant of 3 x 10^, we might choose a diameter of 92 cms.
and length of 40 cms. Eight poles woiild be suitable for a machine of this size.
Although we wish to generate only 250 volts, it will be found that a lap winding
will require about 124 conductors per pole, or 8 x 124=982 conductors in all ; each
carrying 100 amperes. Now we know that 200 amperes per conductor would cut
down the insulation space and give us a cheaper machine. This is possible with a
multiplex winding. Take in this case 2a = 4, because there are 8 poles. We want
about 124 conductors in series, or about 496 -~ 2 = 248 commutator bars. Take the
formula Kfn='(yxp)±a,

and find a suitable y. Try y=61.

ir,„ = (61x4) + 2 = 246.

This number of commutator bars would allow us to have 82 slots with 3 bars
per slot. As the armature stamping is made in one piece on an armature of this
size, 82 slots is permissible. We thus have about 10 slots per pole, a sufficiently
great and yet an economical number.

In the calcidation sheet given on page 514 the machine has been worked out.

We give below enough of the winding table of this machine to show how it goes
and to indicate the bars to which the balancing rings are connected. There are
eleven balancing rings, each connected to two points of the windings. Where

* The reader ia referred to the latter part of the paper by Dr. S. P. Smith and R. S. H. Boulding,
Jaum. LE.E,, vol. 63, p. 232.

W.M. 2 K

DYNAMO-ELECTRIC MACHINERY

.ic^.'Vu/y,.,,/3 T,j^ . cc. cm M

w ?<?e_j Pfi.....: Ph.« -: Volt. ^^O -A.

P. fc.p. p. <«.-. i-tW Amp. p. b,. «m.. 2<?(? . ■

P..P0IH E^ So« /■<.. .

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m leve/ Ci>frvon,\oti 10% 2hr%.

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Fffit-i-nry ^

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b». Toique

s^/i?^'

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Ctfm/7),

•OO18

CONTINUOUS-CURRENT GENERATORS

515

0 = 2, there wiU be two points of equal potential. Beginning at bar 1, we traverse
one-half of the total conductors before we come to a point of the same potential
as bar 1. This is bar 124. Similarly bar 180 is cross-connected to bar 57, which is
just half-way through the total number of steps from bar 180.

Table XX. Winding Table of 200 k.w. Genebatob with Arnold Multiplex

Singly Re-entrant Winding.

2p=8; 2a =4; y=Ql; A:„=246.

The numbers in the Table refer to the numbers of the Commutator Bais. Where
two numbers appear side by side, as 1-124, there is an equalizer connection
between those two numbers.

1-124

123

184

211

148

245

121

182

209

24-147

146

243

119

180-57

207

144

241

117

■ 178

205

142

239

115

176

203-80

140

237

113-236

174

201

138

235

111

172

199

136-13

233

109

170

197

134

231

46-169

107

168

1 195

132

229

105

166

193

69-192

130

227

103

164

1 191

128

225-102

101

162

189

126

223

160

187

124

221

158-35

185

246

122

219

156

183

244

120

217

154

181

242

118

215

91-214

152

179

240

116

213

150

etc.

Table XXI. Showing Arrangement of Equalizing Connections on Arnold
Multiplex Singly Be-entrant Winding. Eleven Equalizer Rings.

Ring No. - - I. ' n.

III.

IV.

VI.

VTT.

VITT.

IX.

XI.

Commutator bar - 1 13

Commutator bar - 1 124 > 136

(

24
147

35
158

46
169

57
180

69
192

80
203

91
214

102
225

113
236

In going through the calculation sheet there are one or two points that arise on
this slow-speed machine. It will be seen that the iron loss is extremely low, on
account of the low frequency. The teeth are worked at B = 21,400, and yet the
iron loss is less than one-quarter of the armature copper loss. It is well that it is so,
because the total surfaces of the armature iron and winding are not able to dissipate
very much more than the 7450 watts lost, and of this 6100 is armature copper loss.

The saturation of the teeth is so high that it is well to calculate Kg (see page 71),
and to correct the apparent flux-density 21,400 to 21,000 by means of Fig. 46.

The copper is worked at only 332 amperes per sq. cm. Even at this current
density, the end windings should be well opened out so that the air can get between
individual coils.

516 DYNAMO-ELECTRIC MACHINERY

THE SPECIFICATION OF C.C. TURBO-GENERATORS.

There is a considerable demand on the market for continuous-current generators
directly connected to steam turbines, particularly where the output is not greater
than 1000 K.w. For larger outputs there is a good deal to be said in favour of
employing an a.c. generator connected to a rotary converter. The loss on the
converter, which may amount to about i per cent., can be saved in the higher
efficiency of the high-speed a.g. generator set. Another solution where a steam
turbine is to be used to generate continuous current is to drive an ordinary slow-
speed generator by means of a double helical gear.

How far the makers of high-speed c.c. generators for direct connection to the
turbines will hold the field will depend upon their success on making thoroughly
reliable generators to run at speeds that are quite suitable for the designers of the
steam turbine. So many successful machines are now running that there appears
to be no doubt that, for small sizes at any rate, the direct-connected generator
will continue to hold the field.

The main difficulties which have occurred in the past with high-speed c.c.
generators are the following :

Changing of the running centre. It has been found almost impossible to build
a machine which would permanently retain its balance with great accuracy. The
insulation on the conductors will always shrink a little, causing sufficient movement
of the conductors to disturb the balance ; so that, however carefully a machine is
built, it will be found that from time to time the balance has altered just a very little,
and the brushes in consequence do not operate well.

Contact between commutator and brushes. It is, of course, important at high
speeds that the commutator shall be perfectly round and run in a true circle, in
order that the carbon brushes may keep in perfectly close contact. It is difficult
to keep a commutator as true as one would like it to be for these high speeds.

Carbon brushes. For a long time metal wire brushes were used to overcome the
difficulty of keeping contact ; but metal brushes cause too great a wear on the copper
of the commutator, and are themselves worn away too fast to give satisfactory
operation. It is now generally conceded that to be entirely satisfactory, a c.c.
generator must be fitted with carbon brushes.

Radial commutator. The plan of employing a commutator, the working faces
of which form planes at right angles to the axis of rotation, has very much simplified
the problem of keeping perfect contact at very high speeds. Any small deficiency
in the balance will not cause the surface of the commutator to throw off the brush.
Certain difficulties were at first encountered in the construction of these radial-faced
commutators, as the expansion and contraction of the metal would sometimes
distort the radial face out of the true plane ; but more recent constructions have
overcome the difficulty, and radial-type conmiutators can now be built to collect
several thousand amperes up to speeds of 3000 r.f.m. ; and even where the want of
balance is quite perceptible on the bearing pedestals, there is not enough motion
at right angles with the face of the brush to interfere with the electrical contact.
The radial conunutator machines are now very widely used for marine work, for which
they are particularly suited, on account of the small amount of attention required.

CONTINUOUS-CURRENT GENERATORS

517

Diameter and length. One difficulty which has been experienced in designing
c.c. turbo-generators of large output and high speed arises from the fact that the
diameter of the armature is limited by mechanical considerations ; and the only
way of increasing the output is by increasing the length. With a great length of
armature iron, the voltage per turn generated in the armature coils becomes so great
that the commutation becomes somewhat sensitive. The importance of having a
low voltage per bar is considered on page 532. In order to overcome this difficulty,
several devices have been employed : one of these is to wind a ring armature so that
the voltage per turn is only one-half what it would be on a drum-wound armature ;
another device is to connect the back of the armature winding to alternate commuta-
tor bars by means of conductors carried between the armature iron and the shaft,
as illustrated in Fig. 438.* A third method is that illustrated in Fig. 435. Here
the armature iron is divided into two sections, each of which may be regarded as an
independent armature of half the length. A main winding, consisting of conductors
of sufficient section to carry the full current, embraces both sections of the iron,
and would by itself constitute a winding having half the desired number of com-
mutator bars per pole. Before this main winding is put into the slots, a number

FiQ. 435. — Auxiliary connectors to intermediate commutator bars. The odd bars 1, 3» 6,
etc., are connected in the ordinary way to the armature winding. The even bars 2, 4, 6. etc.,
are connected to points on the winding by the connectors (shown dotted), which only embrace
the iron of section A.

of auxiliary connectors are placed in the bottom of the slots, and these are connected
to the main winding and to alternate commutator bars in such a way that as we
pass from an odd bar to an even bar through the main winding and the auxiliary
winding, we embrace the iron of only one section of the armature ; and as we pass
from an even bar to an odd bar, we embrace only the other section of the armature
iron. Thus the voltage per bar is only one-half of what it would be if the main
winding were used alone. The advantage of this method over the method shown
in Fig. 438 is that, with it, the self-induction of the auxiliary connectors is given
the same value as the self-induction of the main conductors, and both pass under
the commutating pole in a manner which makes the commutation between odd and
even bars identical with the commutation between even and odd bars. The method,
however, leads to a somewhat more expensive construction than that shown in
Fig. 438 ; and as the latter method will be quite satisfactory under the conditions
obtaining in the machine there illustrated, it has not been thought worth while
to adopt the more expensive construction in that case.

Distance between brush arms. Another difficulty arises in connection with the
distance between brush arms. As the diameter of the commutator is necessarily
restricted, we must have either very few poles or a very short distance between

♦Dr. R. Pohl, "The Development of Turbo-Generatora,'Vottm. LE.E., vol. 40, p, 239.

518 DYNAMO-ELECTRIC MACHINERY

brush arms. If the number of poles is made too few, the current per brush arm
becomes so great that the commutator becomes too long. Hence there has been a
tendency on the part of designers to lower the speed, so as to enable a larger number
of brush arms to be employed on machines of greater output. This diminution of
the speed interferes so much with the efficiency of the steam turbine that the
turbo-set can, under these conditions, no longer compete with a high-speed A.c.
generator connected to a rotary converter. It is believed, however, that up to
sizes of 1000 K.w. at 550 volts satisfactory c.c. turbo-generators can be built,
running at 3000 r.p.m., and that such sets will be not only cheaper but more
efficient than the a.c. generator and rotary converter combination. In much
larger sizes, however, there is no doubt that the brush-arm difficulty prevents
c.c. turbo-generators from being built for very high speeds.

Gompensating winding. A small number of poles on an armature of large output
necessarily entails a large number of ampere-turns per pole ; and it therefore becomes
necessary to provide on these machines a compensating winding which will prevent
undue armature distortion. For small sizes, say up to 300 K.w., successful machines
can be built without a distributed compensation winding, the simple winding of the
commutating pole being sufficient to bring about good commutation. It must not
be thought that the addition of a compensation winding necessarily entails a very
great expense ; because where the armature cross-magnetizing action is completely
neutralized, a very much smaller air-gap can be employed, and the copper in the
shunt winding can therefore be very much reduced. Moreover, every turn that is
put into the compensating winding constitutes a turn on the commutating pole ;
so that the turns adjacent to the commutating pole are correspondingly reduced.

Commutating poles. As the commutating interval is extremely short on these
machines, and the current per brush arm is often very great, the commutating
voltage is often higher than on slow-speed c.c. machines. In cases, however, where
sufficient commutator bars per pole (see page 532) are employed, the commutating
voltage is not too great to be satisfactorily dealt with by means of a commutating
pole and carbon brushes. In fact, it cannot be said that the difficulties met with in
the past have been commutating difficulties; they have rather been difficulties
of collecting the current from a rapidly revolving metal surface.

Critical speed. It is found very difficult on many c.c. turbo-generators to make
the shaft sufficiently stiff to give a critical speed above the running speed. This
is because the bore of the spider on which the commutator is mounted restricts
the diameter of the shaft. It will, therefore, be generally found that high-speed
c.c. generators run above their critical speed. Where, however, the construction
is sufficiently rigid and proper methods of balancing are employed, this leads to no
practical difficulty, and many such machines are giving very excellent service.

Specification. In drawing up a specification for a c.c. turbo-generator, the
purchaser should have in mind the difficulties which have been met with in the
past ; but he should not so word his specification as to restrict the manufacturer
in his methods of overcoming the difficulty. It is sufficient that he should insist
that the machine put forward shall be free from the troubles met with in the past.
It will be seen from the model specification given below in what way this can be
achieved.

CONTINUOUS-CURREIJT GENERATORS 519

SPECIFICATION No. 13.

STEAM TURBINE CONTmUOUS-CURRENT GENERATOR SET.

170. The work covered by this specification is to be carried oenemi
put in accordance with the general conditions and regulations ""^"■^
issued by and dated the day

of 19

171. It includes the supply, delivery, erection, testing, Extent of
finishing and setting to work in the Corporation's Generating ''^*-
Station at of the following plant :

Section I. One high-pressure steam turbine.

Section II. One 1000 k.w. continuous-current generator
direct-connected to the steam turbine.

Section III. One surface condenser with air-pump and
circulating pump of sufficient capacity for the above-men-
tioned turbo set, together with the motors for driving the
same.

Section IV. All pipe work and valve work between the
turbine and the condenser and between the condenser and
its air and circulating pump.

172. The turbo set is to be erected on the site shown in the
accompanying drawing No. , or as may be shown on
additional drawings furnished by the engineer of the Corpora-
tion or supplied by the contractor and approved by the said
engineer.

173. Any fittings, apparatus or accessories which are not Aoceasoriea not
enumerated m this specification, but which are usual or
necessary in the equipment of such plant, are to be provided

by the contractor without extra charge.

174. The contractor is to verify all dimensions and parti-
culars given on the said drawings, and is to obtain all necessary
measurements on site.

175. As the contract for the buildings and foundations has Alternative
been let and the same are in hand, the contractor will be Dra^n^.
required to arrange his plant to suit them. Drawings of

620 DYNAMO-ELECTRIC MACHINERY

the buildings and foundations will be supplied for the use
of the contractor, and may be seen at the offices of the
Corporation for the purposes of the tender.

Work carried 1 76. The followiug work will be carried out by the Corpora-

corporation, tion, and is not included in this contract :

(a) The erection of the power-house, including all
work and materials connected with the floors, and the
final floor surface (except such materials as are a neces-
sary part of the plant).

(6) All work and materials required in connection
with excavation and building of trenches and pits, and
filling in and making good, as well as the cutting away
and making good of walls for pipes, supports, etc.
Cover-plates for trenches will be supplied by the
Corporation, except such cover-plates as form a neces-
sary part of the turbo-generator set.

(c) The instalUng of the circulating water-pipes up
to the flanges of the circulating pumps ; the installing
of the main fresh water supply to the generating
station, together with all meters and pipes up to the
connections of the turbo-generator set, if any.

(d) The installing of the main supply and discharge
pipe valves and the gearing for the same, in connection
with the supply and discharge trenches in the engine-
room basement, but not the necessary pipes and valves
between these trenches and the steam-driven generating
plant included in this specification.

Extent of the 177. The work covered by the first and second sections of
8ea>nd°sections thc Specification includes, in addition to the turbine and gene-

ol Specification. j. n xij. i • • j j_' xi. x i.' j

rator, all littmgs, pipes and connections on the turbine and on
the generator, but does not include any work beyond the inlet
flanges of the turbine high-pressure steam separator, or the
outlet flange of the exhaust steam stop-valve, or beyond the
exhaust flange of the turbine, nor does it include any cable
work or trench work beyond the terminals of the generator.
The work includes the high-pressure steam separator of the
turbine and a field regulating resistance and surface-plate
multiple contact switch for the same.

Extent of the (Here follows a statement of the extent of the work on the condenser.)

Third Section
of Specification.

Extent of (Here follows a statement of the extent of the work on the pipes and

Fourth Section ^^i^pj, px \
of Specification, vaives, etc.;

CONTINUOUS-CURRENT GENERATORS 521

(See Clauses 6, p. 271 ; 36-7, p. 360 ; 74, p. 382 ; 272, p. 591.) Foundations.

(See Clauses 125, p. 461 ; 55-59, p. 379.) Acoeasibiuty

of Site.

(See Clauses 8, p. 271 ; 60, p. 379 ; 273, p. 591.) Use of Crane.

SECTION I.

Turbine.

(This does not fall within the province of this book.)

SECTION II.

Generator.

178. The generator is to be of the compound- wound Rating and
continuous-current type, having a drum armature with the characteristics.
winding in open slots. The machine is to be provided with
compensating windings and commutating poles, and is to
be suitable for supplying current for traction purposes. • It
shall have the characteristics set out below :

Normal output 1000 k.w.

Voltage at full load * 600.

* Where a machine is intended for lighting as well as for traction service, the
generator will be described as a compound and shunt wound generator, and the
voltage characteristics may be stated as follows :

Normal voltage on traction 600 volts.

Compounding From 676 to 600.

Amperes on traction 1670.

Normal voltage on lighting 480 volts.

Amperes on lighting 1700.

Adjustment of voltage on rheo-
stat when machine is run-
ning as a shunt generator 460 to 500 volts.

Regulation as a shimt generator 10 per cent, drop in voltage between no load

and tvdl load, the speed being consteuit
or, and the rheostat untouched ;

Regulation of set, generator

running as shunt machine The speed regulation of the turbine and the

inherent regulation of the generator shall
be such that the voltage shall not fall by
more than 16 per cent, between no load
and full load.

A claiise is sometimes added to the effect that the governor may be altered
when running on traction so as to give 10 per cent, higher speed. This has tiie
advantage that it enables the generator to work at the best state of saturation
both on lighting and on traction.

522 DYNAMO-ELECTRIC MACfflNERY

Amperes. 1670.

Voltage variation, no

load to full load 575 to 600.

Speed Fixed by maker.

Over load 26 per cent, for 1 hour.

50 per cent, for 5 minutes.
Excitation Shunt and series coils.

Temperature rise after
6 hours full load run 46° C. by thermometer.

56° C. by resistance.
Temperature rise after
2 hours 25 per cent.

over load 60° C. by thermometer.

65° C. by resistance.
Puncture test 1500 volts alternating applied

for 1 minute between copper
and iron.

Horuontauy 1 79 . The field frame shall be spUt horizontally and arranged

*^ **' so that the armature can be Ufted out without disconnecting

more than a minimum number of field connections.

Critical Speed. 180. The rcvolviug part of the generator shall be so con-
structed that the critical speed differs from the running speed
by not less than 700 revs, per minute.

Balance. 181. The rcvolviug armature shall be so constructed that

practically no relative motion shall occur between the
constituent parts after completion. The armature shall be
balanced with extreme accuracy so that no perceptible
vibration is communicated to the bearings ; approved
means shall be provided both on the commutator and on the
armature for fixing balance weights and enabling the same
to be readily changed in position.

Noise. 1 82 The generator shall be enclosed so that it shall not give

rise to any more noise than would be observable in machines
of similar size built according to the best practice.

Factor of 183. At thc uormal speed chosen by the maker the calcu-

lated factor of safety in every part shall not be less than 4.
The revolving parts shall, before leaving the contractor's
works, be run at a speed 10 per cent, above normal without
showing any signs of movement of the component parts
relatively to one another.

CONTINUOUS-CURRENT GENERATORS 523

184. The armature is to be provided with a half coupling, coupung.
and is to be driven by a half coupling supplied and fixed to
the end of the turbine shaft. T^sToupESg shall be of an
approved type.

(See Clauses 67, p. 380 ; 268, p. 590.) BearingB.

(See CJlause 68, p. 381.) Eddy Currento

' ^ ' in Shaft.

(See aauses 24a, p. 334 ; 33 and 35, p. 360 ; 68-9, p. 381.) shaft.

1 86. The generator is to be fixed to a bedplate of approved Bedplate,
construction which shall give sufficient rigidity, having
regard to the type of foundations proposed, and ensure the
true alignment of the turbine and armature shafts at aU
times.

186. The holding-down bolts and foundation plates are Hoidingdown
to be provided by the contractor.

Bolts.

187. The armature shall be built up of punchings, which t^ of
shall be supported on the spider or on the shaft in such a
manner that there is no possibihty of their becoming loose or
moving even by the smallest amount relatively to the shaft.
The supports for the end windings shall be such that they can

be readily taken off and replaced in case it should be necessary
to repair the armature windings, and the armature winding
shall be of such a type that a new coil can be inserted without
any great delay.

188. The commutator and brush gear shall be designed so commutator
that there is no tendency for the brushes to be thrown off the ^^
commutator when running at a high speed, notwithstanding

small errors in the balance of the armature.

189. The brushes shall be of carbon. Bnuhea.

190. A drawing of the proposed commutator and brush Drawing of
gear shall be suppUed, showing the method of supporting aubSftt^ ^th
the commutator bars and securing them against centrifugal ^*^°***'*
forces, and showing also the method of making the connections
between the armature conductors and the commutator
segments.

191. The commutator is to be built up of hard-drawn copper conBtruction of
segments accurately spaced circumferentially, insulated from ^JS^^^***'

624

DYNAMO-ELECTRIC MACHINERY

Brush G«ar.

Sample BruBh-
holder.

one another by built-up mica of uniform thickness ; the whole
being supported on mica bushes in such a manner as to avoid
ra,S «/Lal dfapUcement. The arrangement, for ventila-
tion shall be such as to secure a continual supply of cold air
playing over the cooling surfaces of the commutator. Means
are to be provided for tightening-up bolts or nuts on the
commutator centring bushes without the necessity of dis-
turbing any of the armature connections, and also for con-
trolling any movement due to the expansion and contraction
of the segments. The mica bushes separating the commutator
from its metal retaining bushes shaU present a creeping
distance of not less than 1 inch. The tenderer shall state the
depth of copper allowed for wear. Particulars shall be given
of the number and sizes of the brushes proposed and the mean
current density in the brushes with a load of 1670 amperes on
the generator. Provision is to be made for adjusting the
position of the whole set of brushes and arms simultaneously
by means of a worm gear or in other approved manner. The
brush gear is to be designed so that all adjustment and
replacement of brushes and the cleaning of the commutator
can be done when the machine is running.

192. A sample brush-holder shall be suppUed with the
tender.

Commutation. 193. The generator shall be designed to nm at all loads
up to full load under all conditions as to voltage regulation
specified with a fixed position of the brushes, without any
sparking visible at the corners of the brushes.

Throwing Load 194. It shall bc possiblc to throw on full load and throw
off 100 per cent, overload suddenly without causing flashing
over the commutator.

SimUar
Machines in
Operation.

Ventilation.

195. The tenderer shall give in a schedule a Kst of places
where machines of a similar character can be seen in opera-
tion. Preference will be given to the type of machine which
experience has shown to require the minimum attention to
brush gear during normal operation.

196. If the contractor requires cool air drawn from the
outside of the building to cool the machine in question, he
should state the fact when tendering and show on his tender
drawing how the ducts for supplying such air can be con-
veniently arranged.

CONTINUOUS-CURRENT GENERATORS 626

197. The machine shall be so designed that when running Parauei
as a compound generator it will run well in paraUel with ''"''°'°'-
the continuous-current machines at present instaUed in the
Corporation's generating station. The existing generators
consist of the following sets :

(Here follows list of existing sets.)

198. These sets at present run in parallel when over
compounded from 675 volts no load to 600 volts full load.
The series windings on these machines are connected between
the equaliser bar and the positive bus bar. The voltage
between these bars with full load on the station is 1-5 volts.
The contractor must supply all the series resistances and
diverters which may be necessary to make the generator
supplied by him divfde its load eqiaUy with the other gener-

ators in the station. Drawing No shows the position

of the bus bars and equalising bar and the size of cables
proposed for making connection to the turbo-generator.

199. The main terminals shall be designed to take con- Tenninais.
veniently the size of cable proposed by the Corporation, and

they shall be fixed to the frame and insulated in a substantial
manner. The field winding terminals of the generator are
to be entirely separate. Each cable socket shall be pro-
vided with clamping screws and shall be clamped as weU as
sweated, so as to prevent the cable from falling out if by any
accident it become overheated. All terminals of opposite
polarity are to be arranged so that they are either at least
6 inches apart or are provided with insulating screens which
make the shortest arcing distance between them not less than
6 inches.

200. The contractor is to supply the spare parts set out spawB.
in Schedule I., and he is also to state what other spare parts

he recommends, together with their prices.

201 . The contractor is to provide a full outfit of spanners toois.
and special tools necessary for disassembling and assembling
the generator, together with a rack for holding them,

202. The efficiency of the generator shall be calculated Efficiency,
from the separate losses, which shall be measured as follows :

(a) Iron loaSy friction and windage. The generator shall
be run at full speed at no load as a continuous current motor,

626 DYNAMO-ELECTRIC MACHINERY

the pressure at its terminals being 600, with all brushes in
position and adjusted to their working tension. The power
taken to drive the generator under these conditions shall be
taken as the sum of the iron loss, friction and windage.

(6) Copper losses in armature and field. The resistances
of the armature and all field windings, including the com-
pensating winding and commutating linding,^h diverter
if any, and the series winding with additional resistance,
shall be measured by passing a substantial current through
them and observing the voltage drop. The PR losses at
full load shall then be calculated from these resistances, due
allowance being made for the actual temperature rise on load.
The loss in the field rheostat shall also be included.

(c) Brush losses. The brush losses shall be taken as equal
to the watts obtained by multiplying the armature current
by 2 volts. The contractor shall state in his tender the
efficiency, calculated from the separate losses, which he
guarantees at ftdl load, three quarter load and half load ;
he shall also guarantee that the efficiency under actual
running conditions will not be more than 1 per cent, lower
than the efficiencies so calculated.

Testa at 203. Thc foUowiug tcsts shall be carried out at the con-

MaKer s 91*1 o ^ * /*i

Works, tractors works m the presence oi the engineer oi the
Corporation, before being forwarded for erection at the
station :

(a) Iron loss, friction and windage measurement.

(See Clause 1686, p. 503.)

(6) Copper loss measurement.

(See also Clause 235, p. 564.)

(c) Commutation test on short-circuit. The positive and
negative terminals shall be short-circuited through an
amperemeter, the machine run at full speed, and the current
brought up to full-load value and maintained there for
3 hours to test the commutating qualities of the machine.
The machine shall not be forwarded until all adjustments
have been made that shall be necessary to bring about
perfect commutation.

{d) Puncture test. At the conclusion of the short-circuit
tests when the armature is still hot, a puncture test of 1500
volts alternating shall be applied between the armature
winding and frame and between the field winding and frame.

CONTINUOUS-CURRENT GENERATORS 527

204. After erection in the station of the Corporation the gg*^.*'**'
following tests shall be made :

(a) Temperature run. The generator shall be run for
6 hours on full normal load, viz. 1670 amperes at 600 volts,
and shall then be shut down with all possible speed, and the
temperature of the principal parts taken by means of a
thermometer. The temperature of the shunt winding shall
also be calculated from its rise in resistance. No part of
the generator shall rise in temperature more than 46° C.
measured by thermometer, or 50° C. measured by rise of
resistance.

(6) Over-load run. Immediately after taking the tempera-
tures mentioned in the last paragraph, the generator shall
be run at 25 per cent, over load for 2 hours without exhibiting
a rise of temperature of more than 55° C. by thermometer,
of 65° C. by resistance.

(c) Commutation. During the full-load temperature run,
the commutation shall be noted and shaU not be deemed
satisfactory if any sparks are visible at the edges of the
brushes. On 25 per cent, over load there shall not be sufficient
sparking to injure in any way the brushes or the commutator.
After the fall load run there shaU be no apparent marking of
the commutator.

(d) Switching in and out. Full load shall be suddenly
switched on to the generator while it is going at full speed,
and the circuit breaker shall be opened oy having 100 per
cent, over load suddenly thrown on the generator. The
generator shall not flash over or be otherwise injured by this
treatment.

(e) ParaUeling test. The generator shaU be run in paraUel
with any one or a number of the existing generators, and
shall divide the load with them sufficiently well for practical
purposes.

(/) Compounding test. The load shall be varied from no
load to full load to see that the generator compounds from
575 to 600 volts.

{g) Absence of undue noise and vibration. It shall not
be possible to hear the generator running outside the station
of the Corporation, and the vibration of the pedestals and
bedplate shall be not greater than is observable on the
machines of the best construction.

205. In measuring the temperature rise the atmospheric Temperature
temperature shall be taken by means of a thermometer

528

DYNAMO-ELECTRIC MACHINERY

Maintenance
Period.

Provision of
Load and
Steam.

Provision of
Instruments.

placed in the flume or duct which brings the ventilating air
to the base of the machine.

206. The satisfactory completion of the above tests shall
not exonerate the contractor from Uability in connection with
the good running of the plant during the first six months
after the set has been accepted and taken over. If during
this six months defects in the commutation or any other
defects due to faulty construction or design, or bad workman-
ship, shall become apparent, the same shall be immediately
rectified by the contractor, and any time which shall elapse
between the notice given to the contractor of such defect and
the remedying of the same shall not be included in the six
months' maintenance period.

■

207. The Corporation will provide the means of loading
the generator for the temperature tests, and they will supply
all steam and labour for such tests free of charge. They
will also supply free of charge steam and labour equivalent
to 6 hours' full load run to enable the contractor to adjust
the plant. Any additional power which the contractor may
require will be supplied to him at the cost of one hal^enny
per unit. Should the first official tests not be satisfactory,
they are to be repeated, and the cost of such additional
tests shall be borne by the contractor.

208. The contractor shall provide all standard instru-
ments necessary for the foregoing tests and pay for the
calibration of the same.

Cleaning and
Painting.

Drawings
supplied by
the Corpora-
tion.

209. Before delivery all rough parts of the turbo-generator
shall be properly filled, and it shall be given one coat of
paint. After it has been erected and tested in the station,
and when it is ready for continuous working, it shall be
properly cleaned, and painted by the contractor with two
coats of oil paint of approved colour, and one coat of varnish.
If required by the Corporation, this painting may be deferred
until the end of the term of maintenance.

210. Drawing No. supplied with this specification shows
the existing lay-out in the generating station of the Corpora-
tion and the proposed site for the new turbo-generator set.
The contractor is advised to inspect the site and make all
necessary measurements. The contractor is to be responsible

CONTINUOUS-CURRENT GENERATORS 529

for obtaining any information which shall be necessary to
him in deciding as to the suitabihty of the site for his plant
and for the exact dimensions of all foundations, pipes,
flanges and clearances, and other matters with which he may
be concerned.

211. Schedule No. 2 gives a list of the drawings and Drawtop to b©
samples which are to be submitted with the tender. t&. "^^^

212. A provisional sum of £100 is to be included in the Provisional
total price submitted for the whole of the work, which sum ^^'
will be dealt with in accordance with Clause of the
General Conditions.

213. The tender shall state on what date the plant will be Deuvery.
erected and ready for work.

DESIGN OF A 1000-K.V. CONTINUOUS-CURRENT TURBO-GENERATOR

As we have seen, it is desirable to settle upon as high a speed as possible, but this
must at the same time be consistent with thoroughly good performance.

The main difficulties in the way of choosing very high speeds are as follows.
Just as with a.c. generators, the maximum diameter that can be chosen will depend
upon the speed, and the tendency among designers is to employ a somewhat smaller
diameter for c.c. generators than for a.c. generators of the same speed. The
reason is that the winding of a c.c. generator, being subjected to a fairly great
voltage between turns, requires even greater care in its insulation and cannot
be supported quite so well mechanically as the winding of the field magnet of an
A.c. generator. The wedges in the tops of the slots must be made of some insulating
material, such as fibre, instead of brass. Moreover, the connections to the com-
mutator necks are likely to give trouble if too high a peripheral speed is chosen.
A peripheral speed of 15,000 feet per minute, or say 75 metres per second, seems
to be thought very high for o.c. turbo-generators, but if suitable provision is made
to meet the difficulties mentioned, there is no doubt that speeds up to 17,000 feet
per minute are quite possible while still preserving good factors of safety. It is
desirable to keep the diameter as large as possible in c.c. turbo-generators of great
output, because one wishes to keep the axial length as short as possible in order to
get good commutation. It must be remembered that the armature (unlike the
field magnet) must have its core built of laminated iron, and the shaft must not
form part of the magnetic circuit. We will generally find that after we have made
the shaft of sufficient diameter to give it the right stiffness, having regard to the
great span between bearings necessitated by the long commutator, and after we have
provided space for a suitable spider to bring in a sufficient quantity of ventilating
air (see page 206) there is none too much room left for the iron core, so that every

W.M. 2 L

630 DYNAMO-ELECTRIC MACHINERY

quarter of an inch that we can gain radially is going to help us considerably in
shortening the machine.

In the case of the machine under consideration it will be found that a speed of
2750 R.P.M. is not too high for a diameter of 24 inches, or, as we are working in
centimetres, say, 61 cms.

Choice of the number of poles. The number of poles to be chosen depends as
with slow-speed machines upon the current to be collected, but as the speed is
high, and it is not desirable to unduly increase the frequency, one does not choose
any more poles than are necessary, having regard to the current per brush arm.
Thus, on a slow-speed machine one would often choose 400 amperes per brush arm
in preference to 500, while in a high-speed turbo-generator one is often compelled
to choose 1000 amperes per brush arm rather than increase the frequency from
80 cycles to 120 cycles. The distance between the brush arms on the commutator
is abo an important consideration. As the diameter of the commutator is neces-
sarily restricted, we cannot unduly increase the brush arms without getting them
very near together and increasing the danger of a flash-over.

In the case under consideration we have at full load 1670 amperes, so that if
we make a four-pole machine we will have 835 amperes per brush arm. As we know
that many successful machines have been built with over 1000 amperes per brush
arm, we accept four poles as satisfactory in this respect. If the diameter of the
commutator is such that the speed is 50 metres per second (a speed by no means
too high), the pitch distance of the brush arms may be -a^ much as 27 cms. This,
though not as much as we would like, is more than exists on many traction generators
which are working well. If we were to choose six poles, the pitch of the brush arms
would be reduced to 18 cms., rather a short distance for a traction generator, though
quite permissible on a good commutating generator designed to work on a steady
load. Having settled these preliminaries, we can take a calculation sheet and
proceed. The machine is illustrated in Figs. 435 to 439.

The pole arc. This is settled by considering the distance that we would like
to preserve between poles. With a diameter of 61 cms., giving a field bore of say
63 cms., the pole pitch may be taken roughly at 48 cms. Allowing 5 cms. for the
width of the commutating pole, and another 5 cms. for space between iron, we
arrive at the dimension shown in Fig. 436, which gives us a pole arc of 32-5 cms.
It is a good plan to bevel off somewhat the edge of the pole as shown. This makes
the coefficient ^/=0-68. As we are dealing with a continuous-current machine
with a full-pitch winding, we also have ^« = 0-68 (see page 13).

610 =0-68 X 46 X 36 X AgB,

The number of commutator segments. There is no feature in the design of a
traction generator more important than the provision of enough commutator
segments per pole. When slow-speed engines are used to drive traction generators,
it is found that in order to build a machine which will withstand bad short circuits
on the trolley wires without flashing over, it is necessary to have a large number of
commutator bars per pole. It is good practice to have as many as 48 bars, or even
more, per pole. With a design of this kind, the advantage is that even if the load for
an instant is so high that good commutation is impossible, the voltage per bar

CONTINTJOUS-CURRENT GENERATORS

531

Amp* p. cood. 4:2.0 . Amp* p. br. Mrm.O^Q

Wnffw* VnOrVK W^TlllWir ^.r^T... .PolC* tC^.-ElCC- SpCC fA^,..

K.W /OQO; PF ; PhaM . VolU.SSQr.6p/}..; Amp* per ter./57(C?..: Cr^.S^^..; R.P.U^SO.] Rotor Apot

Temp rise .4v$ **.<?.< ReguUtioni!^ Cit7/^f^ Overload ZSjHxtrihmr

Customer : uroer no

;'*««^59 C.rc»«/92:Gm»Are./^/0;^".

:»C?:.(S8 . 7 VoHs=-6ax..-^6 x .36 «

Order No ; Quot- No ; Perf. Spec ; Fly-wheel effect

Ag B ....; pose. UZa

av5%2 X/(0?. I.Z. eo.SQO

; Qxnxm..Sf.S. K. V. A.

''■^"'•SS^-^x/^^

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Ann. A.T. p. pole 7^0..

Max. Fid. A.T.

Armatupe. Rev

Dia
Dia

Outs.

ft.
o
o

Ins

Gross Length

Air Vents — ^

Opening Min >^i0Mean

Air Velocity

Net Lcngtk^A:3 X 89

Depth b Slots

Section _ ^30 .Vol

Flux Density

Loss^l/^p cu.C^^ Total
Buried Cu.2fi^2£L Total
Gap Area-.2<2QC. Wts
Vent Area3P^«0Q<L. Wts
Outs. Area I2J>00, Wts

LOOP
LQ.700

\.7QOO
$300

jAoh

0)
0)

i Noof Segs Ly_|Mn.Circ.
NoofSlotsLZ^|x/-/5=

Section Teeth >-

Volume Teeth

Flux Density.

Loss -/T-p. cu-CffL-Total

c/>

■o
c
o
o

W'eight of Iron.

Star or Mesh Throw

Cond. p Slot

Total Conds 36in SSOSk '.
Size of Cond.4pi y^l'JiS).

Amp. p. sq

Length in Slots -^JL:5_-
Length outside<^jfij&Sum
Total Length lliROL
Wt of i.ooojflZfi^Tofal
Res. p. 1. 000 '/ZS-Total

Watts p.-fll

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Watts p. So

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12 bars pir^ho/e

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Gap Co-eff . K.

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lilL

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Flux per Pole-^'^^/^^' ^

Leakage nl t'O i.lW^JOJxJO^

cage n.L
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Unbalanced Pull I '

No.ofSeg.^^

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A.T. p Pole n.Load S4G0

A.T. p. Polef.Load ^dOO '6707rp €^O0

Surface 2LBQCI ;

Surface p. Wati_ , /<? !

1« R \f3y0 520 /350

I R. ^50 03/ I O'S

Amps. S/O i 900 /67d

No.* of Turns /370 / 3

Mean 1. Turn f ?^~'S t-Sm

Total Lttigth $3QQm. €m i/6fn.

Resistance i22£OkU^2hgL '00038:00043
Res. per i ooo- _ 13- 1 ['Q^7 J^026__
Size of Cond 'Qt^^ACOi 3s^cfh S'S

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Paths in parallel

3fZe ^-Ss^.Z/n

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Magnetization Curve.

Core

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Rotor Teeth
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jQjiQii

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jQ6D^

55jQ. Volts.

A.T.P c«( A.T.

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Field Loss S/lunt

Arm &c TR

Brush Loss

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/20

130(L
1^40

Comnnutator.

Dia. 4iL^peed«5'^
Bars J44

Volts p Rar /7

Brs. p. Arm ^^

Size of Bis. 26^5
Amps p. sq cm. 4 '2
Brusli Loss^SOO+^OC

Watts p. Sq. cyyy 0-4

Mag. Cur
Perm. Stat. Slot
,. Rot. Slot x
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177
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: X. =

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Max. Torque _

Max. H.P

Slip

Power Factor

632 DYNAMO-ELECTRIC MACHINERY

is too low to maintain an arc between successive bars, so that a flash on the com-
mutator occurring at the instant of the short circuit is not readily carried around
from brush to brush. There is a certain critical voltage per bar at which the liability
to flash is very much increased. This critical voltage depends somewhat on the
arrangement of the field system. For an ordinary uncompensated field syst-em
an average voltage of 20 volts per bar is near the danger limit, while for a perfectly
compensated machine the critical voltage per bar may be taken as somewhat higher.
Still, it is good practice even on a compensated machine to keep the volts per bar
well below 20 if it can be done.

If we have 36 bars per pole on the commutator, it will give us 17 volts per bar.
A higher number of bars would give a better performance, but if we take more than
18 coils per pole it will be difficult to find room for them on an armature 61 cms. in
diameter. As mentioned on page 517, the method which we propose to adopt for
obtaining 36 bars per pole when we have only 18 coils per pole is that in which
connectors are brought from the back of the armature to alternate bars. The
method of supporting these connectors is illustrated in Fig. 439. Thus, while we
have 36 bars per pole, we have only 36 conductors in series, so that the armature
ampere-turns are exactly half what they would be if we had 36 complete coils per
pole.

At this point it is well to check the ampere-wires per centimetre. The current
per conductor will be 420 amperes, which multiplied by 144 gives us 60,500 ampere-
wires, and 316 ampere- wires per cm. This figure is not too high ; but by reason
of the small depth of slot necessary, if we are to maintain good commutation at
high frequency, it is quite high enough.

Size of conductors. It will be found that conductors lying in a slot with only
comparatively thin insulation between copper and iron are so well cooled that they
can be worked at a fairly high-current density — ^in this case as high as 430 amperes
per sq. cm. ; whereas in the end connectors, which must be huddled together and
subjected to very much poorer cooling conditions, the current density ought to be
low. It is quite worth while, in a case of this kind, to use a different copper section
for the end connectors, and also a different shape. Electric welding has now been
carried to such a state of perfection that it is a comparatively simple matter to
connect two conductors of different sections by welding. In this case the involute
shape has been chosen for the end connectors, for the following reasons : If the
barrel form of end connector were chosen, the end bell required to support the
connectors would have to project a long way above the surface of the armature ;
and as it is of such a large diameter, the stresses in the end bell due to its own weight
would be exceedingly great. With the involute end connector, supported in the
manner shown in Fig. 438, the supporting rings can be made of great section without
projecting unduly beyond the periphery of the armature, so that the self-stress is
much reduced, and the axial length of the machine is not greater than it would be
for a barrel winding. Another advantage is, that it is comparatively a simple
matter to provide for a wide ventilating duct between the two halves of the end
connector. An end view of the involute connectors can be seen from Fig. 439.
The finger-plates of the armature are made of silicon bronze of great tensile strength
and high conductivity. The two outer end plates are made of the same metal.

CONTINUOUS-CURRENT GENERATORS 633

In between the inner and outer end plates are bridge pieces made of phosphor
bronze castings, which are supported by double spigots on the end plates. The
lower tier of armature conductors, with their end connectors, which are of stranded
copper, are assembled around the armature on a large diameter, with the involute
end connectors properly interleaved. The diameter at which they are originally
assembled is sufficiently great to get over the difficulty of interleaving the con-
nectors. The diameter is then reduced in stages until the lower tier lies in the
bottom of the slots. The outer tier is then assembled in the same manner
and inserted in the slots ; thimble connectors are then used to bridge between
the inner and outer connectors, as shown in Fig. 438. These connectors are
suitably taped over, but allow sufficient room for air to enter the ventilating
duct between the tiers. The ventilating air gets between the bridge pieces where
these flank the conductors ; and at the point where the bridge piece supports
the top conductor, two ducts are milled out to enable the air to escape.

The method of supporting the connectors from the alternate bars to the back of
the armature is also seen in Fig. 439. These are grouped in4)atches, each batch
surrounded by a metal sheath, which ia threaded through a slot in the arm of the
spider. It will be seen that there are 14 connectors in each sheath, although only
twelve are required to go to the commutator bars. The outer connectors of each
batch are short-circuited together so as to form a closed conductor embracing the
air space between the arms of the spider. The object of this short-circuited con-
ductor is to reduce the self-induction of the connectors lying nearest to the air
space. It will be seen that in operation, while one connector is carrying an increasing
current, the connector next to it must be carrying a decreasing current ; so that the
self-induction of the connector can only be due to a flux which lies between itself
and the adjacent conductor. It will be good practice to make the straight part of
the conductors lying in the slots also of stranded copper ; but if solid conductors
are used, it is best to make the conductor near the mouth of the slot shallower than
the conductor at the bottom of the slot. As the frequency is 92 cycles, a solid
conductor near the mouth of the slot would be subjected to serious eddy-currents
unless its depth were reduced to about 11 cms. The conductor at the bottom of
the slot can, however, be made 1 -4 cms. without much fear of eddy-currents (see
page 144) ; so that the average height of a conductor would be 1*26 cms. If the
width is 0-9 cm., the whole will go in a slot 1 16 x 3 cms. The wedge at the top of
the slot is by preference made to the shape shown in the drawing, the depth of the
wedge being 0-7 cm. It wiU be seen from the calculation sheet, page 631, that the
watts per sq. cm. amount to 0 078, which with a thickness of insulation equal to
0 126 cm. will give a difference of temperature between copper and iron of 8° C.

Magnetic loading. A rough calculation^ or our previous experience, leads us to
allow about 10 volts drop in the armature at full load ; so we may take the generated
voltage to be 610. Thus we get the equation :

610 =0-68 X 46 X 36 X AgB,
^^8=0-542x108.

Saturation in the teeth. The mean circle through the teeth ia 174 cms. The
total width of all the slots is 72 x 1 -15 =83 ; so that the total width of all the teeth

536 DYNAMO-ELECTRIC MACfflNERY

is 91 cms. Now, at this high frequency, it is not desirable to work the teeth at a
high density ; so that a length of iron has been chosen to keep the density as low
as 14,600 lines per sq. cm. The gross length is 53-5 cms., and the nett length 41 cms.
91 X 41 gives 3700 sq. cms. section of the teeth. The volume of the teeth is 13,300
cu. cms. ; and as the loss is 0 17 watt per cu. cm., the total loss in the teeth is
2300 watts. The flux-density behind the slots is 10,700, giving a loss of 0-12 watt
per cu. cm., and a total iron loss of 9300.

The investigation of the cooling conditions will be easily followed from the
calculation sheet taken in conjunction with the description given on page 324.
The arrangement of the compensating winding can be easily followed from Fig. 436
and the calculation sheet.

Commutating windings. In working out the strength of the commutating
pole, we must refer to the drawings. Adopting the formulae given on page 480 :

840
Bc= 2-8x41x^ = 1650.

5-00

The axial length of the commutating pole will be about 32 cms. ; we must increase

Be in the ratio ^ to give us 2760. The effective ampere-turns per pole should be

1 X 2760 X 1 08 X 0-796 = 2370.

The armature ampere-turns per pole are 7600 ; so that the total ampere-turns
on the commutating pole, including those on the compensating winding, should be
10,000. We can obtain these ampere-turns by three turns on the compensating
winding and three turns on the commutating pole. It will be found convenient
to divide the compensating winding into two paths in parallel, each carrying half
the current, so that each bar has only to carry half the current. Thus we get
twelve bars per pole, as shown in the drawing. The method of working out the
saturation curve, the cooling conditions on the field winding, and the efficiency, will
be easily followed from the calculation sheet.

Commutator. This is of the radial type provided with 8 grooves, or 16 working
faces. There are 16 brushes per arm, each measuring 2*5x 5 cms., giving a total
area of 400 sq. cms. per terminal, whereas the density is only 4-2 amperes per
sq. cm. A burnt graphite brush of a type similar to the l.f.c. brush of the Le
Carbone Company is suitable for this purpose. The total brush losses amount to
7-5 K.w. This gives us 0-4 watt per sq. cm. estimated on the surface of the com-
mutator ; but as a great part of the heat is conducted into the brushes and brush-
holders, which afford a very well- ventilated cooling surface, the temperature of the
commutator will not be too high. The method adopted of connecting between the

538 DYNAMOELECTRIC MACHINERY

armature conductors and the commutator bars is illustrated in Fig. 438. Special
U-shaped connectors are assembled on the spider in groups, the whole being held
in position by a ring which is divided into two parts and screwed together in the
manner shown. This ring carries a number of projections, upon which the outer
end ring of silicon bronze is supported after the armature winding has been put
into place.

CHAPTER XIX.

ROTARY CONVERTERS.

In what cases they are suitable. For general purposes the rotary converter is
the most efficient machine for converting from alternating to continuous current.
There are some cases in which the motor-generator is more suitable than the rotary
for conversion from A.O. to c.c. If the alternating voltage is very unsteady, and
it is required to have a very steady continuous voltage, then the motor-generator
is the better machine to use. Again, if it is desired to reduce the continuous voltage
to zero and to bring it gradually to full value (as, for instance, in the Ward-Leonard
<5ontrol), the motor-generator is the machine generally specified, though it would be
possible to design a modified rotary converter for this class of work. It used to
be said that a motor-generator had the advantage of being more easily started after
a general shut-down, but now that rotaries are made self-starting and self-synchroniz-
ing, the contention no longer holds.

Many of the large continuous-current railway and tramway systems of the
world obtain their current through rotary converters, and a frequency of 25 cycles
has been conamonly used for such systems. During the last ten or twelve years
the 50-cycle rotary has become established as a perfectly reliable machine both
for traction and general lighting and power supply. There is no doubt that where
a new system is being put down, mainly for continuous-current traction, 25 cycles
or 33 cycles will be chosen in preference to 50 cycles. But where a system of
50 cycles is already in use, there is no difficulty in supplying continuous current for
tramway purposes by means of rotary converters suitably designed to meet severe
conditions sometimes occurring in traction service.

The best frequency. In general it may be said that for fairly high-voltage
continuous-current work (750 to 1500 volts) low frequency is to be preferred, because
it allows us to design the rotary with a greater distance between brush arms than
would be possible with a high frequency. For low- voltage work, as, for instance,
in the supply of continuous current at 250 volts or lower for electrolytic purposes,
the higher frequency is suitable, because it gives a cheaper machine with a large
number of brush arms, and the current per brush arm is not so high as it would be
on a low-frequency machine.

The pitch of the brushes. The distance between the brush arms and the fre-
quency are related on account of the following consideration : During one complete

540

DYNAMO-ELECTRIC MACHINERY

cycle, a point on the commutator must travel through a distance equal to twice
the distance between two consecutive brush arms. Thus, in a SO-cycle converter, if
the pitch of the brushes is 10 ins., then the speed of the commutator must be 20 ins.
in one-fiftieth of a second, or 1000 ins. (about 25 metres) per second ; that is,
5000 feet per minute. As it is thought that this is' about as high a speed as one
should run an ordinary commutator, 10 ins. is about the maximum pitch one will
find on 50-cycle converters. For the same commutator speed the brush arms could
have a 20-inch pitch on a 25-cycle rotary, though for reasons of economy the pitch
is more commonly about 13 ins. It is quite likely that, as the design of brush
holders and the construction of commutators is improved, the commutator speed
will be increased and the distance between brush arms may be increased where a
greater distance is desired.

Variation of the voltage. The differences in the characteristics of various
installations of rotary converters lies mainly in the range of voltage over which the
machines are designed to work and the different methods * employed to effect the
change in the voltage at which the continuous current is supplied.

Putting out of account for the moment the split-pole converter and other machinea
of a similar nature, we may say that the ratio between the voltage on the slip-rings
and the voltage on the commutator remains nearly constant f independently
of the excitation.

Table XXII. gives the values of the ratios between the voltage on the slip-rings
and the voltage on the brushes on the commutator for two-phase, three-phase
and six-phase machines as ordinarily built. These ratios ordinarily have not in
practice the values they would have if the field-form of the machine were sinusoidal.
The widening of the pole (as commonly done on rotaries to increase the output)
has the effect of making the ratio of the a.c. voltage to the c.c. voltage somewhat
smaller than it would be with a sine-wave field-form.

Table XXII.

Ratios op a.c. to c.c. Voltage on Rotaky Converters as apfected by the Ratio
OF Polk Abo to Poiif Pitch, aixowing for Normal Bevelling of Poles.

Pole arc
Pole pitch

Single-phase -

Three-phase -

Four-phase -

Six-phase rings 1 and 4
„ rings 1 and 3
,, rings 1 and 2

0-8

0-76

0-7

0-65

0-6

0-68

0-696

0-71

0-72

0-74

0-6

0-61

0-62

0-635

0-65

0-48

0-49

0-5

0-62

0-53

0-68

0-695

0-71

0-72

0-74

0-6

0-61

0-62

0-635

0-66

0-34

0-36

0-356

0-365

0-376

* See page 546.

t We say " nearly constant," because it is possible by greatly over-exciting or under-exciting
a three-phase rotary, to change the voltage on the commutator some 3 per cent., while the voltage
on the slip-rings remains constant. On a three-phase machine there is at certain instants a
considerable angle between the connection to a slip-ring and the connection to the commutator
bars upon which the brushes momentarily touch. The leading or lagging current passing through
this part of the winding will give a positive or negative boosting effect, but this effect is not
sufficient to give a wide range of voltage.

ROTARY CONVERTERS

541

Table XXIII.
Ratios of a.c. Amperes per Slip-Bino to c.c. Amperes per Terminal,

ASSUMING AN EfFICODBNCY OF 06 PER CeNT.

Pole aro
Pole pitch

Single-phase
Three-phase
Four-phase
Six-phase

0-8

1-54
103
0-77
0-616

0-75

1-51
101
0-74
0-5

0-7

1-47
0-98
0-73
0-40

0*65

1-44
0-94
0-72
0-48

0-6

1-4
0-93
0-69
0-46

Table No. XXIII. gives the values of the ratios of a.c. amperes per slip-ring to
c.c. amperes per terminal, assuming an efficiency of 96 per cent.

These values are given on the assumption that brushes are placed on the neutral
point, the excitation normal, and that the field-form is such as one finds on ordinary
commercial machines.

O-i

0-1

«—

- -

- . .^

foti

on C

1C.6

fenei

rata

oc- -

-- •>

WtntHnq between Tuppinqs

Fia. 500. — Showing the heating of the various parts of the winding of a 3-phase rotary
oonverter running on unity power factor compared with the heathig of the conductors of a
0.0. armature of the same output. Asi-04 ; jfe^o.

The heating of the annatnre conductors. The theory of the heating of the
armature conductors of a rotary converter is so fully dealt with in standard

542

DYNAMO-ELECTRIC MACHINERY

text-books ♦ that there is no necessity to give it here. The heating is greater in
those conductors lying near the points from which tappings are taken than in the
conductors lying midway between tappings.

Figs. 500 and 501 show the rate of production of heat in the various parts of the
armature winding of a 3- phase rotary converter as compared with the rate of pro-
duction of heat in the same winding used as a o.o. generator. The curves in this>

7-8

i /5

::toa

§
C

^04

03
02
01

"¥-

- - ■*>

^ " •

"" \

1
1

1
1
1

1 1

i
1

eati

^n^i

7fa

C.Gt

'ner

ator

1
1

{

:>^

.si

6§

...^

Winding between Tappings

Fig. 501. — Showing the heating of the various parts of the winding of a 3-phaae rotary
converter with current leading by IS"" as compared with the heating of a o.o. generator carrying
the same load. A = 1*04 ; Jt= 0*26.

figure and in the two following figures have been plotted on the assumption that
the power supplied on the a.c. side of the converter is 4 per cent, greater than the
power given out on the c.o. side. At unity power factor (Fig. 500) the heating
is least at the point in the winding midway between the tappings, and it is here
only 0*25 of the heating on a c.c. generator. At points close to the t«ppingB>
however, the heating is greater than on a c.c. generator.

♦ Barr and Archibald, The Design of AUenuUing-Current Machinery (Whittaker), 1913. Wood-
bridge, Proc. Amer^ I.E.E,, vol. xxvii. p. 204.

ROTARY CONVERTERS

543

If m is the number of slip-rings and a is the angular distance of any point P
from the mid-point M of a section of the winding, and 4> is the angle of lag of the
current, the heating of the winding at the point P, expressed as a fraction of the
o.c. heating, is

1 +

•2

Trmsm

(1)

In this formula Hx =JW+W, where h is the ratio of a.c. power to c.c. power
and k is the ratio of the wattless current to the power current at unity efficiency.
Thus, where 4 per cent, losses are supplied by the A.c. current, A =1*04, and where
the wattless current is 0-26 of the working current (<^ =15®), then k =0-26.

^ 1:1

I 10

^ OS

gOS

••4

J 06

'S 0-3
S

•S

"^ 01

HeatirtQ of C.C.GeriercUor

Winding bstwmerv Tappings

Fia. 502. — Showing the heatins of the various
parts of the winding of a d-pnase converter
running at unity power factor as compared with
the heating of a o.c. generator carrying the same
load. ft=104: *=0.

1. 1

1* 1

1-0

1 1 1 1 1

Heatirtq of CC.Generatc

o>d

\r'0

yfTi

J-

—

es.i

0 1

O-S

(VA

)

0-2
O'l

S m^b

n — ^

•

•■

V •

Li_

Windinjg between. Tapping

FiQ. 508. — Showing the heating of the various
parts of the winding of a S-phase converter with
the current leading oy 15** as compared with the
heating of a o.o. generator carrying the same
load. A =104; Jt=0'26.

Fig. 501 shows the efiect of making the current lead by 15® in a three-phase
winding. The winding is supposed to be moving from right to left. In front of
each tapping the winding tends to get hot, the heating effect in some parts of
the copper being 1*8 times as great as in a c.c. machine carrying the same load.
When the current lags, it is the winding immediately following the tapping that
gets hot.

In a six-phase machine the heating is much more evenly distributed over the
winding. Fig. 502 shows the heating at unity power factor, and Fig. 503 the
heating where the current leads by 15°.

It will be seen that at unity power factor, the loss in the copper near the
tapping is 043 of the c.c. loss, while at power factor 0*966 the loss is 0-74 of
the c.c. loss.

544

DYNAMO-ELECTRIC MACfflNERY

The average loss in the whole winding is dependent upon the power factor.
The ratio of the loss in the converter to the loss in the c.o. generator at the same
load is

,^_sH^_m ^2)

.2 •

m^sin*—
m

For a six-phase machine m=6, and if A=l the above expression simplifies

down to

0-268 + 0-89*2 (3)

The second term of this expression gives us the loss due to the wattless load.
As a matter of fact, owing to eddy currents, the loss is greater than given by this
expression (see page 144).

Sf '02 S3 -€4 'S5 'S6 '87 -Sg '69

1-3

J-Z

^^^

/•/

i-a

HilTi

f oi

C, (rtneralOi

'<L

iTe—

•7

■^

5*s

. 1

•o
•5

•3

2jff

SHi

b=:

•2

■

Pon

faci

'vr

Si -92 '93 '94 '95 -96 '97 -99

Fio. 504. — Average loss in the annature conductors of a rotary converter as compared with
tibe loss in the same winding used as a o.c. generator.

If we neglect the eddy-current losses and take A =1, the heating coefficients for
3-phase, 4-phase and 6-phase rotary converters vary with the power factor in
the manner shown in Fig. 604. These curves serve to calculate the efficiency
where the efficiency is arrived at by the measurement of separate losses, but the
actual loss will be somewhat heavier (see page 545).

Now the question arises, how much extra copper are we to put in the arnaatuie
when the power factor is less than unity, say 0-966 ? Consider the six-phase case.
We are not merely concerned with the raising of the average loss from 0-27 to
0-33 in the case given above. Nor are we to say that the temperature will be raised
in proportion to the losses in the conductor adjacent to the tapping (from 0-43 to
0-74 in the case given), because the heat is rapidly conducted away from a single hot
conductor contiguous with others. We are really most concerned with the rise in

ROTARY CONVERTERS 546

temperature of the group of conductors lying near the tapping, and we may take the
following rule as giving results which are accurate enough for practical purposes.

For a six-phase machine, take the working current (at unity power factor) for
any load as if it caused a loss equal to 0*3 of the loss there would be in a c.c. generator
with the same load. Take the wattless current as if it were entirely independent,
and calculate the loss just as we would for a c.c. generator with the same current.
This is the same as taking the coefficient of the second term in (3), page 544, as
equal to unity. Now add the two losses, and the sum will be nearly proportional
to the rise in temperature of the hottest part.

Example 59. On a 6-phase conyerter. Powerfactor 0*966, leading current 0-28 of working
current. Take working current loss as 0-3 and wattless current loss as (0-28)^ =0-078.
Loss =0*378, the loss there would be in a c.c. dynamo carrying the same current. Note : this
method does not apply when calculating the total loss in the armature for the purpose of getting
the efficiency. For this latter purpose the figure would be 0-33 as given above, not 0*378.

Example 60. In a case of a 1000 e.w. converter we are required to supply 300
wattless K.v.A. on the h.t. side of the transformer when running at full load. Before we can
begin to get leading current on the h.t. side of the transformer, we must supply its magnetizing
current from the converter and also have the current leading enough on the low-tension side
to neutralize the tendency of the h.t. current to lag, owing to the self-induction of the transformer.

The transformer in this case will be built for small reactance, so an allowance of a further 100
leading wattless e.v.a. will be sufficient to cover the difference in phase between the low-tension
and the high-tension currents. Strictly speaking, we should get this information from the
transformer designer, but in practice, in fixing these preliminary matters, one makes an allowance
and proceeds. The exact effect of the transformer self-induction will be seen later (see Fig. 595).
We then want 400 wattless K. v.A. on the low- tension side, that is to say, 04 of full-load working-
current leading. This affects the design of the converter in two particulars. It not only calls
for more copper in the armature ; it calls for more copper in the field magnet in order that the
converter may be sufficiently over-excited. For the moment we are only concerned with the
armature copper. The working current is 1900 amperes, and on a 14-pole machine will be
136 amperes per conductor. Take the leading current at 0*4 of this, or 55-4 amperes. Through
one ohm of resistance the working current would produce

136 X 136 X 1 X 0-3 =5550 watts loss.

While the wattless current would produce

55-4 X 55*4 X 1 =3000 watts loss.

Total 8550 watts. While on a continuous-current generator one ohm of resistance would give
136 X 136 = 18,500 watts, so the heating is 8550/18500=0-46 of what it would be in a continuous-
current generator. Now the size of the copper strap required depends upon the number that
are grouped together in one coil and the cooling conditions on the surface of the coil. If our
«ight conductors per slot are insulated with mica and paper and tape to a thickness of 0-05 inch
and say 0-01 inch of air space, it is easy, from the considerations given on page 222, to show that
for a difference in temperature between copper and iron of 20° G. we can pass at least 0-8 watt
per sq. inch. Now our coil will have a surfiace of about 36 sq. ins. per foot run, so that it will
pass to the iron about 29 watts per foot run. If we have a barrel winding fairly well ventilated,
-we will at a high s}>eed easily get rid of more than 0-8 watt per sq. in. in the part outside the
slots. In calculating the loss in the slots we must not forget the eddy-current loss in the straps
on the top of the slot. For straps ^ inch deep and with the dimensions of slot given in Fig. 513,
it will be found by calculation like that given on page 149 that the eddy-current loss is 0-4 of
the loss calculated from current and resistance only. This gives the coefficient 1-4 used
"below. If now we take the efficiency at 90 per cent, and the current to be converted 142 amperes
per conductor, we have the resistance of the strap per foot length at the running temperature

x X 8 X 142 X 142 X 0*46 x 1 -4=29 watts per foot run,
2;=000028 ohm per foot &t 65''C.,
X =0-00024 ohm per foot at 25° C.
w.M. 2 M

646 DYNAMO-ELECTRIC MACHINERY

The sectioo must be such as to have a rwistance of 0-24 (rfun per 1000 ft.

Area of strap -q- sT" =0-034 sq. m.

This would give us an apparent current density of 4200 amperes per sq. inch. Bat before
deciding on this size we most look at the over-load gnaranteesL Soppose that there is an
over load of 25 per cent, for three hoars. This, so £sr as heating is concerned, is the same as if
it were a continaoos over load, becaase a high-speed machine like this will get very near
its maximnm temperatare in the armatare copper in the coarse of an hour's ran. It may be,
however, that there is no gaarantee to ran on a leading power factor at 25 per cent, over load,
and it may be that there is no reason why it should not be ran at or near unity power factor.
Now it is easy to show by the rule given above that the heating on 25 per cent, over load at unity
power factor is about the same as the heating at full load with the addition of 300 wattless k. v.a.
In one case we have 176x175x0 3 =9200,

and in the other 142 x 142 x 0-46 =9300.

If we are allowed 50° C. rise on 25 per cent, over load, no* special provision need be made to
meet this guarantee. There is sometimes the question whether or not it is worth while to put
a little more copper in the armature so that the same design will do for a machine which has
to run under more stringent conditions. That question is generally settled by the standard
size of copper straps kept in stock and the room available in the slot for which we happen to have
a die. We cannot take account of these things here.

In actual practice, of course, we do not go through this long calculation. We know from
experience that we can on a six-phase converter work the copper at from 3000 amperes per sq. in.
to 6000 amperes per sq. in. according to the power factor. We choose a suitable standard strap
accordingly, and if in doubt we check the cooling on the basis of 0-8 watt per sq. in. for a 500 volt
insulation.

Variation of voltage. The voltage can be changed (1) by rocking the bmahes,
(2) by changing the field-form. Either of these methods can be used to vary the
voltage on the continuous-current side, while the a.c. voltage remains constant ;
but the more usual practice, when it is required to change the c.c. voltage, is to
change the a.c. voltage supplied to the taps on the armature winding.

This can be effected by one of the following methods :

(1) By changing the excitation of the a.c. generator supplying the rotary.
This can be done when the a.c. generator supplies nothing but the rotary load,
and it is desired to change at the same time the voltage of all the rotaries connected
to the generator ; for instance, where turbo-generators and rotaries are put down
for supplying continuous current, it may be for electrolytic work. Here a very
wide range of voltage variation can be obtained.

(2) By means of a synchronous a.c. booster in series with the conductors supply-
ing the rotary. A booster of this kind is usually mounted on the shaft of the
rotary, and has the same number of poles, so that it is necessarily synchronous.
The advantage of this method of changing the voltage is, that it gives complete
control of the voltage of each rotary independently, and at the same time allows of
an independent adjustment of the power factor.

The armature may be placed between the slip-rings and the taps to the rotary
winding, as shown in Fig. 516, or it may be outside the slip-rings and connected to
the low-tension side between the transformer and the slip-rings. When the current
to be dealt with is very great, and the voltage of a number of rotaries must be varied
at the same time, it may be convenient to connect an external booster in series
with the high-tension side of the transformer. In this case very great care should

ROTARY CONVERTERS 547

be taken with the insulation of the booster, or the factor of safety of the whole
plant will be lowered. Sometimea it may be convenient to drive the a.c. booster
by means of a synchronous motor, but this plan is not so efficient oi ao desirable
as connecting it to the rotary itself. The design of an a.o. booster ia worked out
on page 582. The right way of making the connections between the armature
wiadings of the booater and the armature windings of the rotary are given in Fig. 506.
A suitable diagram of connections for a rotary and booster is given in Fig. 506.

coDVErter wlndingg vhlch are Inmub, uidatso Uw i>osltian ol the polea.

(3) A third way of changing the A.c, voltage supphed to a rotary is by means of
an inductance in the circuit between the rotary and the supply mains. If a lagging
current is drawn through this inductance, it gives a drop in voltage at the sHp-
ringa. If a leading current is drawn, it raises the voltage. The lagging or leading
current is obtained by under-exciting or over-exciting the rotary. The self-induction
may be obtained by suitably designing the transformer feeding the rotary, or special
reactance coils may be added in aeries with either the high-tension mains or the
low-tendon mains. The most usual method is to design the transformer ao as to
have considerable self-induction, because the losses are very little it at all increased
by so doing, whereas reactance coils would add considerably to the PR and iron

548

DYNAMO-ELECTRIC MACHINERY

losses. Where the reactance required is not too great (10 or 12 per cent.), it can be
obtained by simply grouping the high-tension coils of the transformer together
and the low-tension coils together, so as to cause magnetic leakage between them.
This somewhat cheapens the transformer and increases its factor of safety. Where

Uftfiauuf XwuuLWf Uiujujuf

ivmraiL^^^

Amfdf

Fio. 506. — Diagram of conDectiona of rotary converter and booster to auxiliary switches.

a very high reactance is required, it may be necessary to add iron to the leakage
paths in the transformer. One advantage of external reactance coils is that they
can be readily cut in and out of service as required.

The considerations which determine the amount of reactance to have in
circuit are reviewed on page 599, where the theory of the added reactance is
worked out.

(4) A fourth way of varying the voltage is by means of an induction regulator.
This in effect is the same as an a.c. booster, but an induction regulator is more
expensive than a rotating booster, and is not self-ventilating. It has, however.

ROTARY CONVERTERS 649

the advantage that it does not interfere with the commutation, and may, therefore,
be used for a very wide range of voltage.

(5) A fifth way is by means of taps on the transformer. These taps may be
either on the high-tension side of the transformer or the low-tension side. In
the case of large transformers, they are usually put on the high-tension side, because
on the low-tension side the cost of bringing out taps is much greater, and the voltage
per turn is usually too high a percentage of the whole voltage.

The main difficulty with this method is in the changing of the taps when on
load. Unless some device of a more or less complicated nature is added, we have
a sudden jump in the voltage as we pass from one tap to the next. Moreover, it
is necessary to connect to one tap before disconnecting from the last, and this causes
a short circuit on part of the transformer windings lying between the taps, unless
a " preventative " resistance or some other equivalent device is employed. Modem
designs of the controllers for connecting two successive taps, and for the prevention
of sparking, have made this method much more acceptable than it has been in the
past. Where a really satisfactory means of changing the taps can be employed,
this method of changing the voltage can be recommended. It is more efficient
than any other method, and for large units and wide ranges its first cost is
lower.

One way of stopping sparking at the controller is to employ a small booster
or induction regulator to graduaUy boost the pressure derived from one tap until
it is the same as that of the next tap above. The connection can then be made
without danger of short circuiting, and one can pass from tap to tap with a gradually
increasing or decreasing voltage. For big instaUations, where the expense of such
an arrangement is justified, there is no doubt that its higher efficiency will commend
it. It is possible to arrange the mechanism so that the whole range of voltage is
obtained automatically by the mere turning of a handle.

Changerover switches. It is sometimes advantageous to employ taps on the
transformer for giving fairly wide ranges of voltage variation, and a booster to give
the intermediate voltage values. Where it is desired to give a voltage variation from
410 to 490 volts for lighting, and from 525 to 575 for traction, and where it is
important to preserve a good power factor at all voltages, it is a good plan to arrange
tappings on the transformer (see Fig. 506), so that on one tapping one can get,
without any boosting, 450 volts, the booster being arranged to boost down 10 per
cent, and up to 10 per cent., to give the full range from 410 to 490 ; and another
tapping giving a normal voltage of 550 without boosting, the booster being employed
for obtaining the necessary compounding action. It will generally be found more
convenient to arrange the tappings on the high-tension side of the transformer
than on the low-tension side ; and where it is not intended to make frequent changes
from lighting to traction, ordinary isolating switches are sufficiently convenient
for making the change-over. Where, however, it is desired to change-over fre-
quently, inter-connected oil switches can be employed which will throw out one
tapping and throw in the other without shutting down the machine. Where the
Rosenberg method of starting rotary converters is employed (see page 557), it will
be found .quite easy to switch over from one voltage to another without shutting
down the machine ; all that is necessary is to switch o£E from the lighting bus-bars.

650 DYNAMO-ELECTRIC MACHINERY

put in tlie starting motor, throw over the change-over switch on the high-tension
side, and then parallel the machine on the traction bus-bars.

Regulation. When the load on the continuous-current side of an ordinary
shunt-wound rotary is suddenly changed, the drop in the voltage which will occur
depends upon the amount of drop in the line, the reactive drop in the transformer
and the resistance drop in the transformer and rotary. If we maintain the high-
tension voltage steady in the sub-station and have a transformer with fairly small
reactance drop on full load (cos^=0), the voltage regulation will be extremely
good (assuming always that the brushes are on the netural point). Only the ohmic
drop will then affect the voltage, and this is generally from 2^ to 1| per cent, in
500- volt machines of 200 k.w. to 1000 k.w. capacity. For machines of low voltage,
the ohmic drop is rather greater, on account of the greater drop on the brushes.
If the brushes are rocked forward, the drop in voltage will be greater ; and if they
are rocked backward, it will be less. If there is more reactance in the transformer
(say up to 20 per cent.), the drop in voltage on load is somewhat greater, and the
machine behaves very like a good shunt-wound c.c. generator. If it is desired to
have a very considerable drop od full load, a series coil should be added to the
field magnet, connected in such a way as to weaken the field as the load comes on.
This is sometimes called a reversed series coil. If then there is some reactance
in the transformer, the lagging current drawn from the line causes a drop in the
transformer, and the voltage may be made to fall 10 or 20 per cent, as desired
(see page 595). A rotary with a series coil connected in this way tends to maintain
its load at a steady value, notwithstanding changes in the load of machines running
in parallel. A reversed series coil, then, gives stability of load.

Compounding. When it is desired to make the voltage rise as the Ipad comes
on, the series coil is connected so as to strengthen the field, and the reactance in
the transformer then gives a boosting effect. We give below (page 595) the method
of working out quantitatively the voltage rise with a given amount of reactance
and over-excitation. Where a rotary is fitted with commutating poles, the effect
of rocking the brushes is very mtich more marked than on a machine without
commutating poles. Eveu without a series winding on the main poles, it is possible to
obtain 2 or 3 per cent, over-compounding on the commutating poles only by rocking
the brushes back. This method of compounding is not permissible where the rotary
has to run in parallel with other rotaries or with c.c. machines, because, there
being no equalizer, the load would be unstable, and as soon as the rotary took any
load it would tend to take more and more, and probably bring out the circuit-
breaker on over load. If, on the other hand, the rotary by chance took current
from the line, running as a motor, it would tend to take more and more current
to bring into operation the reverse-current mechanism of the circuit-breaker. We
get the most stable conditions of running by having the brushes rocked forward
of the neutral point. The more we rock them back, the better we make the regula-
tion and the more unstable we make the conditions of nmning. The rocking of the
brushes is the most convenient method of obtaining a fine adjustment of the com-
poimding, and of making a machine share its load with other machines in parallel.

When a rotary is fitted with an a.c. booster, the compounding can be effected
by putting series coils on the field magnet of the booster, through which the main

ROTARY CONVERTERS

661

contmuous current will paes and Btrengthen the field of the booster sa the load
comee on. Sometimes it is an advantage to have series coik on both the main poles
of the rotary and on the booBter. The effect is then to maintain the power factor
as the load comes on, and at the same time to obtain an additional boosting effect.

caoo*

r jnj'...... — ^

. (fiviiiMriiv Diary.)

Bqualtxing. Where a rotary is over-compounded and it is intended to run
in parallel with other compound-wound rotaries or o.c. generators, it is necessary
to have an equalizing bus-bar just as with c.C. generators. The resistances of
the equalizer bar and connections should be kept as low as possible, the object
in view being to feed all aeries coils iirom a common source and at one common
voltage. Any voltage drop in an equalizer comiection which tends to give a machine

662 DYNAMO-ELECTRIC MACfflNERY

a higher voltage on its series winding than exists on the other series windings, will
tend to cause instability. Each machine has (particularly if its brushes are rocked
forward) a certain amount of inherent stability. Now, the amount of instability
caused to any machine by the resistance of the equalizer cable must not be greater
than the amount of inherent stability possessed by that machine. Fig. 507 shows the
method of connecting to the equalizer bar.

Dependence of G.G. voltage on A.G. voltage. Where the a.c. voltage fluctuates,
the c.c. voltage will also fluctuate, the percentage variation being the same on both
sides of the machine. In cases where it is desired to maintain the c.c. voltage
steady in spite of variations in the a.c. voltage, an automatic regulator may be
arranged to control the field of an a.c. booster so as to compensate for the variations ;
but where the variations on the a.c. side are very sudden, and it is of great import-
ance to keep the c.c. voltage very steady, it is better to use a motor-generator, as
in this there is no electrical or magnetic connection between the c.c. side and the
A.c. side, and as long as the speed remains constant, the voltage generated is not
aflected by variations of the a.c. supply.

Three-wire machine. Where a rotary feeds a 3-wire lighting network, it is possible
to make it act as a balancer by merely connecting the mid-wire to the star-point
of the low-tension side of the transformers. The connections are shown in Fig. 507.
In normal cases, the resistances of the rotary and transformer being very small,
this balancing effect is exceedingly efficient. A rotary will run with one side fully
loaded and the other side unloaded, the whole return current going to the star-
point of the transformer. The commutation under these conditions will be quite
good, and the drop in voltage on the loaded side need not be more than 3 per cent.
Where it is desired to use a rotary as a balancer, and at the same time to compensate
for ohmic drop in the mid-wire, a booster may be connected in circuit between the
mid-wire and the star-point of the transformer. This mid-wire booster may be
either series wound or may carry fine wire winding, the current through which can
be controlled by hand.

Power factor. The rotary converter being a synchronous motor on the a.c.
side, its power factor can be adjusted by adjusting the excitation of its field-magnet.
Its efficiency is highest when running at unity power factor, but it is oft^n
desired to make it take a leading current to compensate for lagging currents taken
by other apparatus on the system. When it is intended to call for a rotary for this
purpose it is better to specify the amount of wattless leading current required
than to specify the power factor, for while the leading current may remain constant
the power factor will change with the load. Moreover, specifying the amount of
leading current required warns the man who draws the specification what he i»
asking for. To call for a 1000 k.w. rotary converter that shall run on 90 per cent.
leading power factor at full load seems a reasonable request, and does not on the
face of it appear to involve a much greater expense than to call for a 1000 K.w.
rotary to run on unity power factor. But when we remember that this means a
machine which must, in addition to its 1000 K.w. load, yield 484 k.v.a. wattless,
and that this wattless k.v.a. by itself would cause nearly twice as much heating
of the windings as the true kilowatt load, we pause to consider whether as much
wattless load is really required. Another matter which needs careful consideration

ROTARY CONVERTERS 553

when calling for wattless leading current is the amount of self-induction that must
be put in the transformer to meet other requirements in the specification. Suppose
that we call for a rotary without booster that is to be over-compounded 10 per cent,
between no load and full load. As will be seen later (see page 596), this will involve
the use of a transformer with sufficient self-induction in it to give a reactive drop
of about 20 per cent., and the rotary will have to yield a leading current equal
to about 0-3 of its full-load current, in order to produce the required boosting effect.
This leading current flows between the rotary and the low-tension winding of the
transformer. But on the high-tension side of the transformer, the phase of the
E.M.F. being different by reason of the self-induction, the current is almost in phase
with the E.M.F. That is to say, that while the rotary is yielding a big leading
current which is tending to heat it up, this current is not available for compensating
for lagging currents in the high-tension system. It is, in fact, absorbed in correcting
the power factor of the leaky transformer. Now, if we were to call upon this plant
to yield 0-3 of full-load working current leading, in addition to its other load, it
would mean that the rotary armature would have to supply a wattless current
equal to 0-6 of the full-load working current, and the heating effect would be three
times as great as on the same machine (6-phase) working on unity power factor.

Where the converter is required to yield a leading current, and at the same time
to have a variable voltage, it is good practice to adopt one of the methods for
voltage variation (such as an a.c. booster) which permits the transformer to be
built with only a small self-induction. The armature is then not called upon to
supply much more than the leading current furnished to the system.

Parallel miming. The troubles from hunting which used to occur on the early
rotary-converters have been overcome by fitting the poles with well designed
dampers or amortisseurs, and by properly adjusting the fly-wheel effect and short-
circuit current to prevent resonance with irregularities in the frequency of supply.

Where the angular speed of the generators supplying the power is perfectly
uniform, as with turbo-generators, no special precautions are necessary beyond
the fitting of suitable dampers to the poles ; but where there is a considerable
fluctuation in the angular speed in the generators, the amount of the fluctuation
and the frequency of the swing should be known, and the fly-wheel effect and short-
circuit current of the rotary adjusted to such values that the unsteadiness is not
aggravated by resonance. On page 337 the laws governing such matters are given ;
and on page 345 we have worked out an example to show how resonance may be
avoided.

The effect of the damper in reducing the phase-swing is considered on page 601.

Starting. Various methods are used for starting rotaries.

(1) Starting on CO, side. When continuous current is always available in the
sub-station where the converter is placed, it is common practice to start up just
as one would start a continuous-current motor, a starting resistance being put in
circuit at first and gradually cut out as the converter comes up to speed. The
speed of the rotary is adjusted by means of the field rheostat. Fig. 506 shows how the
connections may be made for a rotary to be started up on the c.c. side. The method
of connecting the synchroscope is shown in Fig. 508. According to this diagram,
the synchronizing is done on the high-tension side. It may be that at the moment

554 DYNAMO-ELECTBIC MACHINERY

of synchroiuBm the A.o. voltage from the rotary transformer las not the same
virtual value as the a.c. voltage of the maius. If the a.c. switch were closed uiidei
these circumstances, the rotary would immediately take load, the amount of which
would depend on the diffeience between the two a.c. voltages at the instant of

Fio. SOS.— 'Dlasram of eoanectloiu ot e-phue T0UT7 oonTertei lor tnpiilyliia pawn tiom
CO. bi i.O. The macMoe it Bbf>rt«d up fiam the D.c. bare mid lynchranlieif on uia bt. ddB □[
the tmufoimei. The apecd Is ngulated b; the dliBct-conaccMd eiclt«i. (fiivlnarruv Diary)

switching and the regulating quality of the rotary. Under certain conditions, this
load might be excessive ; so it is good practice, when starting up a rotary on the
c.c. side, to open the c.c. circuit-breaker immediately before closing the A.c.
switch. Where the o.c. circuit -breaker is fitted with a trip coil, it is easy to arrange
for the handle which closes the a.c. switch to make in passing (juat an instant
iiefore it closes) a contact which brings out the 0.0. breaker.

ROTARY CONVERTERS 566

(2) Starting by means of a starting motor and synchronizing by hand. This method
is very commonly used on rotaries of large capacity, and is generally regarded as a
thoroughly satisfactory method. The starting motor is commonly an induction
motor with a high-resistance squirrel-cage rotor. The number of pairs of poles is
made one less than on the rotary, so that its synchronous speed is higher than
that of the rotary, and the resistance of the rotor is arranged to give the required
sHp, so that the speed of the rotary may be just right when it is fully excited. This
enables the A.o. switch to be closed without any shock to the system, as connection
is made to an already excited machine. The speed is adjusted by changing the
excitation of the main field. This changes the iron loss of the rotary, and hence
the amount of slip of the motor. It is permissible to close the a.c. switch when the
voltage of the rotary differs by 10 or 15 per cent, from the voltage of the bus-bars,
because the machine is disconnected on the c.c. side and nothing happens except
the flow of some wattless current. On very large converters loading coils are some-
times provided to change the speed ; these are connected to slip-rings and act
simply by putting an a.c. load on one of the phases. If preferred, the induction
motor may be provided with a wound rotor and slip-rings, and a rheostat used
to change the speed.

(3) Starting from taps on the transformer, self-synchronizing . It sometimes happens
that the a.c. supply to the rotary converters feeding a traction system is cut off for a
short time, and all machines are stopped at the ?ame time. When the a.c. supply
comes on again, it is necessary to start up the rotaries as quickly as possible and put
them into service. Now, it may be that in times of stress such as this the a.c. voltage
and frequency are very unsteady ; so that just at the time when it is desirable to syn-
chronize quickly it is most difficult to do so. A method of bringing up the rotary to
speed quickly and throwing it on the bars without waiting is therefore of the greatest
importance. One way of doing this, which is suitable for small rotaries, is by throw-
ing on to the collector rings a voltage of ^ or ^ of the normal voltage, and bringing
the rotary up to speed as an induction motor. The dampers on the poles of the
rotary in this case act like the squirrel cage in a rotor, and give a very considerable
starting torque. The ^ or ^ voltage is obtained by taps on the low-tension side
of the transformer ; so that even if the machine takes three times full-load current
in the armature, the current in the high-tension line has only about full-load value.
As the rotary gets near to synchronous speed, the slip is so small that the voltage
observed on the brushes alternates very slowly. The wattless currents in the
armature magnetize the poles. At first the m.m.f. alternates quickly, but as the
rotary comes up to speed the alternation may be so slow that during the time
that the salient poles are magnetized with a certain polarity the armature may be
dragged into synchronism. When the rotary is up to synchronous speed, the fact
is indicated by the c.c. voltmeter, which then gives a steady reading of J or J full
voltage.

A diagram of connections for this system of starting is shown in Fig. 507.

If the polarity of the brushes is correct (as indicated by the polarized voltmeter
reading on the right side of the scale), the slip-rings of the rotary can be connected
by means of a throw-over switch to a higher voltage, and ultimately to the full
voltage of supply. If the machine comes into synchronism with the wrong polarity,

656 DYNAMO-ELECTRIC MACHINERY

it is necessarj to make it slip a pole before throwing it on to the higher tape. This
majr be done by reveising the connections to the shunt field (by means of a reversiog
switch shown in F^. 607). This makea the B.H.r. of the aimature oppose the
cunent in the shunt coils, so that the cuirent sinks to zeto. As you watch the

c.c. voltmet«r, you see the needle awing to zero as the field dies and the a
b^ina to slip. If, now, the reversing switch be thrown over t^ain (so as to come into
its normal position), just as the armature has slipped one pole, the field will excite
again ; but it is now found that the polarity is right.

This method of starting a rotary is the simplest, and would be satisfactofy
but for two drawbacks : (1) the lai^e wattless current taken &om the line, and (8)

ROTARY CONVERTERS 657

the sparking which occurs under the brushes during starting. This sparking may
be very troublesome, and even where it is slight it may be sufficient to prevent the
commutator from assuming that high state of polish which is so desirable. For
this reason some firms provide apparatus attached to the brush gear to raise the
brushes at starting. For small machines this is not generally regarded as necessary,
and as the current drawn from the line is not so excessive, the method is very
widely used for small rotaries up to say 350 K.w. capacity.

Sometimes two sets of tappings on the low-tension side of the transformer are
provided, so that the voltage applied to the rings can be brought up on easy stages
and the rush of current which occurs on throwing over to the higher tappings is
reduced.

Fig. 507 shows the arrangement for starting a small three-phase rotary
by means of two double-throw triple-pole switches connected to taps on the
transformer.

(4) Starting motor connected in series with slip-rings. Dr. E. Rosenberg has in-
troduced an ingenious method of starting and synchronizing rotaries which possesses
the advantage of rapid self-synchronizing without the disadvantage of taking
heavy wattless currents from the line or causing sparking at the brushes. The
method will be understood from Fig. 509, which gives the connections for a three-
phase rotary. A three-phase starting motor has the six ends of its star winding
brought out. The ends which would be ordinarily starred are connected to three
of the rings of the rotary. The impedance of the winding of the rotary armature
is so low that for practical purposes these ends may be regarded as star connected
at the moment of starting. The other three ends of the motor winding are carried
to the terminals of the transformer A, E, and C. The three-pole low-tension knife
switch shown in the figure is open during the starting. In order to start, the motor
switch is closed and the motor brings the rotary rapidly up to speed, taking only
about 30 per cent, of full-load current. As the voltage across the rings at starting
is quite low, perhaps 6 per cent, of normal voltage, there is no sparking at the
brushes. As the rotary gets up speed it excites itself, but as long as the frequency
of its alternating voltage is different from that of the supply the motor still exerts
a turning moment, though this will be less or greater according as the voltage of
the rings is in or out of phase with the supply voltage. The motor is wound with
one pair of poles less than the rotary, so that it would take it above synchronism
if it were not for the fact that the current through the starting motor provides
enough torque on the rotary acting as a synchronous motor to prevent it from
exceeding the synchronous speed. The condition of synchronism is indicated by
a steady reading on the central-zero voltmeter.

If the starting current is kept fairly low, s^y 30 per cent, of full-load current,
the residual magnetism of the rotary will not be disturbed, and the polarity of the
terminals will be right when the machine gets into step. If it is desired to start
the rotary in a very short time, say 20 seconds, a somewhat larger starting current
must be used, and then to ensure the rotary having the correct polarity when going
into step a field switch is provided which is kept open until synchronous speed has
been almost reached. By watching the voltmeter which moves a little to the right
and left as the rotary slips pole after pole, a moment can be chosen for closing the

558

DYNAMO-ELECTRIC MACHINERY

field switch so that the polarity will come up in the right way. Fig. 510 shows
the arrangements for a six-phase rotary converter.

High-tension tuppiy

High-tension
oil switch

Low-tension "/"/^
knife switch A J A

Step-down br^uisformer

6-phd.se
rotary converter

Central-zero
volt^^er

lOntlnuQiJs-cunent
suppfji

Fio. 610. — Diftgram of connections for st-arting 6-phase oonyerter and self-synchioiiizing

(Rosenberg's method).

Running C.C. to A.G. Sometimes it is required to take power from continuous
current bus-bars and convert it into a.c. power for transmission to some distant
point. The rotary converter is very widely used for this purpose on account of
its high efficiency. Where the converter employed for this purpose has to run
in parallel with a.c. generators of definite frequency, no special precautions
need be taken to regulate the speed of the converter, because its speed will be
synchronous with that of the a.c. generators. Where, however, there is no

ROTARY CONVERTERS 659

generator of definite frequency in parallel with the converter, it is necessary to
regulate the speed and so fix the frequency of supply.

One difficulty in fixing the speed of the converter when it is running without
any synchronous generator in parallel arises from the fact that any wattless load
(current lagging) tends to weaken the field magnet of the converter, and thus to
increase the speed. An increase of speed makes the current lag more, and thus
one gets a cumulative effect that may result in the converter running away.

In order to obviate this difficulty, Mr. B. G. Lamme proposed the use of an
under-saturated exciter driven by the converter. The exciter for this purpose
is generally mounted on the end of the shaft of the converter, and its armature is
electrically connected through a rheostat to the shunt winding of the converter.
The exciter is so designed that at the voltage at which it ordinarily works the iron
of the magnetic circuit is not saturated, that is to say, the normal-voltage point
is below the knee of the saturation curve. Under these circumstances, a slight
increase of speed of the converter makes a very considerable increase in the exciting
current. It is, in fact, possible under practical conditions to obtain an increase
of 5 per cent, in the exciting current for an increase of only 1 per cent, in the speed.
This arrangement is found to work very well, so that even with a varying load of
low power factor, the speed of the converter can be kept within sufficiently narrow
limits. Where heavy over loads of low power factor are expected, it is well to
put series coUs on the converter arranged so that when a load comes on the field
of the converter is strengthened. This plan is sometimes adopted in small con-
verters instead of using the direct-driven exciter. It is effective in those cases
where the wattful load increases at the same time as the wattless load.

After these observations upon matters affecting the operation of rotary con-
verters in general, we can proceed to give moddl specifications such as might
be issued by the engineer of the intending purchaser.

We shall consider two cases : First, a 1250 K.w. 6-phase converter for lighting
and power supply at 460 to 500 volts, as well as for traction work at 525 to 560 volts.
This machine will be fitted with an a.c. booster, so that we can consider the
characteristics of such a generator to work out its design in detail. Secondly,
a 2000 K.W., 250-volt, 6-phase rotary designed for electroljrtic work.

We shall then give some notes on the methods to be adopted to meet special
requirements.

660 DYNAMO-ELECTRIC MACHINERY

SPECIFICATION No. 14.

1260 K.W. ROTARY CONVERTER AND A.C. BOOSTER
(General Clauses Nos. 1, p. 269 ; 21, p. 333 ; 170, p. 519.)

w?rfc*°' 214. This specification provides for the supply, erection,

testing and setting to work of a rotary converter and a.c.
booster set, having the following characteristics :

Characteristics of Rotary Converter.

S'SJJJSteV.'* 216. Normal output :

Running a.c. to c.c.

1260 K.w.

Running c.c. to a.c.

1260 K.v.A. at 0-9 power factor.

Number of phases

Frequency

60 cycles per second.

Continuous-current

' ' . -

voltage

460 to 500 volts for lighting

bus-bars.

626 to 560 volts for traction

bars.

Continuous-current

amperes

2360 amperes.

Three-wire network

Rotary converter to act as a

balancer. Out of balance

current 600 amperes.

Compounding

On traction 626 to 660.

Adjustment of voltage

on rheostat :

A.c. to c.c.

On lighting from 460 to 500

while H.T. volts vary from

6000 to 6400.

c.c. to A.c.

6400 to 6600 while the lighting

•

bufi-bars vary from 460 to 500.

Range of variation of

A.c. voltage

In practice this may vary be-

tween 6200 and 6700, but
performance is only asked for
on the basis of 6400 to 6600.

ROTARY CONVERTERS 561

Leading idle power re-
quired from rotary
and transformer when
running a.c. to c.c.
at full load 490 to
650 volts 300 K.V.A.

Over load 26 per cent, for 3 hours at unity

power factor.
60 per cent, for 10 minutes at
unity power factor.
Temperature rise after
6 hours full load at
560 volts, 0-97
power factor on h.t.
side 40° C. by thermometer.

Temperature rise after
3 hours 25 per cent,
over load 60° C.

Characteristics of Booster.

216a. The a.c. booster shall have a revolving armature characterisucg
situated between the slip-rings and the converter armature. °
Its capacity shall be sufficient to enable the continuous-
current voltage of the converter to be changed gradually at
any load from 460 to 560 when the high-tension voltage
in the sub-station has any value between 6400 and 6600. The
change of voltage may be effected by boosting down in the
lower part of the range and boosting up in the upper part of
the range ; but in proceeding step by step from the lowest
voltage to the highest, there must be a continuous action of
the controlling handle.

216. During the whole range of voltage, the power factor Power-factor
of the set on the high-tension side shall be under control,* so ^^^'
that it can be adjusted at any voltage between the limits of
unity and 0*97 leading, by altering the main field rheostat.

* Where it is not necessary to control the power factor, a somewhat cheaper arrange-
ment may be suitable. In this case, the clause as to power-factor control might read
as follows :

It is not insisted that the power factor of the set shall be completely under control
throughout the whole range of voltage ; it may be lagging (not less than 0*0 at full
load) between the voltages of 460 and 480, and may be leading between the voltages
of 630 and 650. Between the voltages 480 and 530, which are the normal voltages
on lighting and traction respectively, it shall be possible to maintain the power factor
at unity, and at the higher voltages it may be leading.

w.M. 2 N

562

DYNAMO-ELECTRIC MACHINERY

Duty of Plant. 217. Two lotaiy converters of the above rating, with their
boosters and transformers, are required to run in aU existing
power-station for feeding into the low-tension network in
the vicinity of the power-station, and also to form a link
between the existing a.c. generators and c.c. generators.
The generators consist at present of three 3000 K.w. three-
phase steam-turbine-driven sets running at 1500 B.P.M.,
and two 2000 k.w. continuous-current generators driven bv
direct-connected steam engines running at 120 r.p.m. Nor-
mally, one of the 1260-K.w. rotary converters will feed the
lighting bus-bars, and for this purpose taps must be provided
on the high-tension side of the transformer to enable 480
volts c.c. to be obtained at unity power factor without any
appreciable boosting. The other will run normally on the
traction bars, and taps must be provided on the transformer,
which give (at unity power factor) 630 volts without any
boosting.

Interchaiige-
ability.

Bniminff
luverted.

218. Both sets are to be completely interchangeable.

219. On Sundays, or at such other times as it may be
desired to shut down some or all of the A.c. turbo-generators,
the two converters are to be capable of running inverted from
the c.c. generators in the station and of supplying a.c. power
through their transformers at from 6300 to 6600 volts to
outlying districts. It may be that one or two rotaries will
have to supply the whole of the A.c. power ; or it may
be that one or both will have to run in parallel with one
or two of the A.c. generators, assisting in the supply of
A.c. power.

Maintain
Frequency.

220. When running as the only source of a.c. power, they
shall maintain the frequency within 5 per cent, of 50 cycles
per second, provided the wattless k.v.a. does not amount to
more than 300.

Variation of
Load.

221. When running on the traction bus-bars, the current
may vary from 1000 amperes to 3000 amperes from minute to
minute ; the machine must therefore be of liberal design and
commutate well during the peaks of the load.

Bun well In
Parallel on
0.0. Side.

222. It must run well in parallel with one or two of the
present 2000 K.w. c.c. machines, which will be compound-
wound, when running on the traction bus-bars.

ROTARY CONVERTERS 663

223, The voltage drop in the series windings and c^^'J^^j^p
nections at full load on these machines is 2* 5 volts. The "^ ^ ® •
Contractor must provide all diverters and series resistances
necessary to make the rotary converters divide their load
reasonably well with the c.c. generators. The compounding
of the present generators is from 625 volts at no load to
650 at full load.

224. When running in the lighting bus-bars, the rotary con- characteristics

o o ^ o ' • • of Shunt

verters must have the characteristics of shunt machines with Machines.
9 per cent, drop in voltage between no load and full load.

225. The converter must be capable of acting as a balancer, Balancer.
and of dealing with an out-of-balance current of 600 amperes

in the middle wire. With this current flowing, the difference
in voltage between the two sides of the three-wire system shall
not be more than 1 per cent, of the voltage across the outers.

226. When running on Ught load during some parts of the ^^p^
day the rotaries are to be over-excited and to take a leading
current from the line. When running at quarter load,
measured on the c.c. side, each machine must be able to
supply 500 K.v.A. wattless ; and when running on full load
each must be able to supply 300 k.v.a. wattless for six hours
without exceeding 40° C. rise.

227. The voltage of the high-tension bus-bars varies change hi:H.T.
between 6300 and 6600 volts. The design of the rotary ^'^^^^^
converters must be such that this change of pressure will

not cause them to take such excessive loads as to bring out
the circuit breakers or cause trouble from bad commutation.

228. Under all the conditions set out above, the con- stabiuty in
verters must be very stable and free from hunting. operation.

229. The rotary converters must run without any appreci- commutation.
able sparking or glowing of the brushes while the load is
changed from zero to 25 per cent, over load, with the brushes

in a fixed position.

230. Each converter may be started up by means of a Emergency
starting motor or other means, but the arrangements shall ^**'**°*"
be such that in case of emergency it can be switched in on

the A.c. side in less than 1^ minutes from the time of starting
from rest, however unsteady the frequency and voltage may

664

DYNAMO-ELECTRIC MACHINERY

Normal
Starting.

Vibration and
Noise.

be, and even if the voltage is only 90 per cent, of its normal
value. When switched on it must be of right polarity and
inmiediately available for throwing on the c.c. bus-bars,
provided always that the c.c. voltage is high enough. In
the case of this emergency starting, it will be permissible to
draw from the line a momentary current equal to 1-6 times
full-load current, but there must be no necessity to wait for
any indication on any instrument or any synchronizing which
is not perfectly automatic.

231 . Preference will be given to methods of starting which,
while complying with the above requirements for starting on
an emergency, can be so ordered under normal conditions
that there is no shock to the system on throwing in a machine.
For this purpose the machine may be synchronized either
by band or automatically, and the time taken may be de-
pendent upon the steadiness of the frequency and voltage.

232. The sets must run smoothlv and without vibration
under all conditions of load. They must produce no more
noise than is made by machines of similar size and speed
constructed according to the best practice in this respect.

Insulation
Tests.

233. The armature windings and field windings of the
rotary converters are to be subjected to a pressure test of
2000 volts alternating, applied between the windings and
frame for one minute while the machine is hot.

Puncture Test
onlSlte.

234. This test shall be carried out at the maker's works
in the presence of the representative of the Purchaser, and it
shall be repeated after the plant is installed if, in the opinion
of the Purchaser, there is reason to beheve that the windings
have been damaged.

Efficiency. 236. The efficiency of the rotary converter shall be cal-

culated from the separate losses, which shall be measured in
the following way :

(a) Iron loss, friction and windage. The machine shall
be run as a c.c. motor at full speed and at various voltages,
and measurements made of the c.c. power taken to dnve
it, and of the exciting current taken at various voltages.
These tests shall be taken both with the booster fully
excited and with the booster unexcited.

ROTARY CONVERTERS 666

(6) Copper losses. The resistance of the armature and
series field coils of the rotary converter and booster shall
be taken by measuring the voltage drop in them when a
substantial current is passed through them at a known
temperature. The resistance on full load shall be taken
as 1-2 times the resistance at 20° C. The loss in the
converter armature resistance at 0-96 power factor shall
be taken to be 0-378 times the loss on the armature when
working as a continuous-current generator. The resist-
ances of the series coils and commutating winding shall
be taken with any diverters that may be necessary,
suitably attached and adjusted.

(c) The losses in the shunt windings of converter and
booster and their rheostats shall be taken to be the
exciting current multiplied by the voltage of excitation
in each case respectively, except that where any potentio-
meter-type rheostat is employed, the total current going
to the rheostat and field shall be taken as the exciting
current.

(rf) The brush contact losses on the commutator shall
be calculated by multiplying the measured combined
pressure drops at the positive and negative brushes by
the continuous current. The brush contact losses on
the slip rings shall be calculated by multiplying the
measured pressure drop from brush holder support to
shp rings by the current per ring and the number of rings.

236. The Contractor shall state what efficiency * he is ouarantee ot
prepared to guarantee at full load, three-quarter load andEffldencr.
half load, the efficiency being calculated as stated in the last
clause.

Or (see Clause 250, page 686).

Or.

237. The Contractor shall state what efficiency he is Guarantee of
prepared to guarantee at full load, three-quarter load andE^iency.
half load, such efficiency to be measured by means of both
indicating and integrating instnmients on the a.c. and c.c.

sides.

238. The instnmients used in the measurement of the calibration of
efficiency specified in Clause 237 shall be calibrated both "*™™®'*

* Very cominoEily, the converter and its transformers are supplied under the same
contract, and in that case it is usual for the Contractor to give figures for the overall
efficiency of the whole set.

566

DYNAMO-ELECTRIC MACHINERY

before and after the test by some institution, to be agreed
upon by the Contractor and Purchaser.

g^wonof 239. All instruments required for the measurements

"jdPower for aforcsaid shall be provided by the Contractor, and all power
required for one preliminary test and one final test shall be
provided by the Purchaser, free of charge. If either party
shall require a test to be repeated, the party so calling for a
repetition shall pay for the power consumed, unless it shall
appear that he was justified in calling for a new test and that
the necessity for it was not due to his fault. Power for
additional tests shall be supplied by the Purchaser at the
rate of per unit.

Tests.

Terminals and
Connection.

Bearings.
Brush Gear.

Oscillator.

Insulation.

Sample CoU.

Qrinding Gear.
Painting.
Cables.
Spare Parts.

Dates of
Completion.

Drawings.
Fomidations.
Use of Crane.

Accessibility
of Site.

(See Clause 284, p. 592.)

(See Clauses 44, p. 361 ; 112-113, p. 443 ; 199, p. 525 ; 267, p. 590.)

(See Clauses 67, p. 380 ; 268, p. 590.)

(See Clauses 106, p. 442 ; 188, p. 523 ; 191, p. 524 ; 263, p. 689 ; 305,
p. 609; 312, p. 611.)

(See Clause 264, p. 589.)

(See Clauses 93, p. 439 ; 269, p. 590.)

(See Clause 270, p. 590.)

(See Clause 271, p. 591.)

(See Clauses 209, p. 528 ; 278, p. 591.)

(See Clauses 6, p. 271 ; 42, p. 361 ; 73, p. 382 ; 279, p. 591 ; 320,

p. 611.)
(See Clauses 20, p. 274 ; 114, p. 444 ; 169, p. 503 ; 200, p. 525 ; 280,

p. 592.)
(See Clause 281, p. 592.)

(See Clauses 174, p. 519 ; 282-283, p. 592.)

(See Clauses 6, p. 271 ; 36-37, p. 360 ; 74, p. 382 ; 272, p. 591.)
(See Clauses 8, p. 271 ; 60, p. 379 ; 273, p. 591.)
(See Clauses 8, p. 271 ; 55-59, p. 379 ; 125, p. 461.)

Screw threads. (See Clause 277, p. 591.)

ROTARY CONVERTERS 567

DESIGN OF A 1260 K.W., SO-CYCLE, 6-PHASE ROTARY CONVERTER

TO COMPLY WITH SPECIFICATION NO. 14.

Voltage variation. In designing any rotary converter, the first question to
consider is the method by which the variation of voltage is to be carried out, because
the power factor will depend upon the method we employ (see page 546).

In this particular case the voltage is to be varied from 460 to 500 volts on a
lighting load, and from 525 to 550 on a traction load. If there were no objection
to operating the set at a low lagging power factor on the low voltages, and a leading
power factor on the higher voltages, the variation of voltage might be carried out
by means of an inductance in the transformer, in the manner described on page
595. But in this case the purchaser requires the converter to 3deld 300 leading
wattless K.v.A. at all loads, so that it is not permissible to run the converter on a
lagging power factor. The most suitable method, therefore, for obtaining the
voltage variation is by means of an a.c. booster, which by preference shoidd
be mounted on the shaft of the converter between the slip rings and the*
armature.

Having adopted the booster method of voltage vaiiation, the only necessity
for wattless current will be the meeting of the guarantee to deliver 300 leading
wattless K.v.A. on the high-tension side of the transformer.

Taking into account the magnetizing current of the transformer, which in this
case would be supplied by the converter and would amount to about 5 per cent,
of the full-load current, the power factor on the low-tension side of the converter
would be 0-96 leading. If the efficiency of the converter at full load be 96 per
cent., the k.v.a. input would be about 1360. Referring to the curve in Fig. 504,
we see that the loss in the armature conductors will be 0-33 of the loss in an equiva-
lent c.c. generator, but taking all the factors into account which are considered on
pages 544 and 545, we know that the temperature rise of the winding will be about
0-378 of the temperature rise of an equivalent c.c. generator.

Under these conditions, experience shows us that we may take a L^l constant
of about 2 X 10^ (see page 570).

Number of poles. When we are designing a 500-volt rotary converter intended
for traction work, our main aim will be to produce a machine having highly satis-
factory commutating qualities even when carrying a heavy over-load. For this
reason, the number of kilowatts per pole shoidd be made much smaller than woidd
be permissible in a converter intended for a steady load. A rating of 100 k.w.
per pole may be taken as a fairly conservating figure, and some designers might
prefer only 80 K.w. per pole.

In the old days designers were cautious, and knowing that the commutation
was easier when the current per brush arm was small, they built their converters
(particularly the high-frequency converters) with a large number of poles. This
gave very large machines of slow speed, and though the performance was fairly
satisfactory the efficiency was much lower than it need be and the cost was high.
It was soon found that with proper adjustments much larger currents could be
satisfactorily dealt with on each brush arm, so the number of poles was decreased,
the speed increased, with the result that we have now very much cheaper and

668 DYNAMO-ELECTRIC MACHINERY

more efficient machines. How far this reduction in the number of poles will go
in the future it is difficult to say. It is quite possible to build a 1500 K.W., 550-volt,
6-phase, 50-cycle rotary converter, having only 4 poles, and running at 1500
B.P.M., but such a machine would not be cheaper or more efficient than a siz-pole
machine running at 1000 r.p.m. It is doubtful whether the six-pole machine would
be an improvement upon the slower speed machines with which we are more familiar ;
the commutator would be very long, and we should have conditions such as we have
in continuous-current turbo-generators, instead of the easier conditions of engine-
type machines. The saving in cost as we reduce the number of poles is not as great
as we might at first suppose. The conmiutator is one of the most costly parts of a
converter, and we gain nothing in economy by reducing the diameter, for we have
to make it longer in proportion (or even in a greater ratio) if we have to deal with
the same current. The bars on a long commutator, moreover, are deeper than those
on a short conmiutator of the same peripheral speed, so the cost of the commutator
is really increased as the number of poles is reduced. We do not get so much
benefit on a long commutator from the blowing of the conmiutator necks. Now
there is nothing to be saved on the collector gear by increasing the speed of the
converter, for we have the same current to collect. As we have to provide for a
certain cooling surface on each ring for each watt lost, the conditions are only made
more difficult with increased speed. The considerations as to commutator brush
gear and collector are of great importance, because in a sense these are the most
essential features in a rotary converter. We can imagine the other parts of the
machine being done away with.

Weight of copper on the armature. For the same peripheral speed a greater
weight of copper is required when the poles are few than when they are many.
The reason is, that we have in any case the same volume of current to carry from the
slip-rings to the commutator, but when the number of poles is greater the number
of paths in parallel is greater, and therefore the current per slot is smaller, which
enables us to work at a rather higher current density.

We do, however, make a saving in the iron of the magnetic circuits and in the
size of the frame by reducing the number of poles. The relative calculated costs
of commutators, armature windings, armature punchings and frames will be very
different in different factories. In getting out costs, so much depends upon the
rating of the tools used and the apportionment of general charges. It is therefore
impossible to give any rule for arriving at the best number of poles for a rotary
converter of given output.

In the example worked out below, we have taken 14 poles. This number is
suitable, judging from the general practice of to-day.

It gives a rating of 90 k.w. per pole. As the machine is rated at 2360 amperes,
we have 2360-^7=336 amperes per brush-arm at normal load, and 500 amperes
per brush-arm at 50 per cent, over load. These are suitable values for a 50-cycle
converter subjected to heavy fluctuating loads. On a 25-cycle converter we
would work at a higher current per brush-arm in order to reduce the number
of poles and increase the speed.

On low-voltage machines, where the current to be generated is very great, one
is guided more by amperes per brush-arm than by kilowatts per pole. We might.

ROTARY CONVERTERS 569

for instance, take 1000 amperes per brush-arm as a maximum beyond which,
it is not desirable to go, and fix the number of poles accordingly. But on
a 500-volt machine the current per brush-arm is usually kept lower than this.

Diameter and length* These will generally depend upon a manufacturer's
standard sizes ; but if they are to be settled from first principles they would be
controlled by the choice of a suitable pole pitch. This, in a 50-cycle converter,
may conveniently lie between 30 and 36 cms., and really depends upon the room
required for the requisite copper and iron, without exceeding an axial length which
has been found in practice to give good commutation. If we choose a pole pitch
of 32 cms. in this case, the circumference of the armature will be 32 x 14 =448 cms*
Taking a diameter of armature 142 cms., we will find that we can get in the requisite
copper and iron without making the axial length greater than 31 cms. ; so that
this diameter is suitable. These dimensions make the output coefficient =2*1 x 10^.
Before we can calculate Ke, we must settle upon the pole arc. The main con-
sideration in settling this is to leave enough room for the commutating pole and
space between the commutating pole and main pole. On 50-cycle converters, the
space required for these generally amounts to about 25 per cent, of the pole pitch ;
so that the pole arc ought not to exceed 75 per cent, of the pole pitch. In this
case our pole arc is 23-5 cms. ; and as the tips of the poles have a slight bevel, the
coefficient Ke (see page 13) equals 0*74.

Number of bars per pole. A machine of large output of this kind will be in-
variably wound* with a lap winding, there being one turn per commutator bar.
The number of bars will be settled from considerations similar to those given on
page 532, and we may take 48 bars per pole as an entirely satisfactory number ;
so that we shall have 96 conductors in series between the positive and negative
brushes. Allowing 15 volts drop in the resistances of transformer and converter,
we arrive at the formula :

565 volts =0-74 x 7*15 revs, per sec. x 96 x AgQ x 10"®,

AgB^l'llxKfi.

We can now make a general check calculation to see the relation of the AgB

and the I^Za to the size of the frame. The calculation sheet is given on page

570. The circumference = 446 cms., and the area of the working face ^4^= 13,800

sq. cms.

Ill X 10* -^ 13,800 =8000 O.G.S. lines in the air-gap,

168 X 1344 -r 446 =500 ampere- wires per cm. of periphery.

These values are suitable. The length of armature is now settled by the amount
of iron required in the teeth. Figs. 511 to 516 give sectional views of the machine.
Fig. 513 shows the size of slot and the arrangement of the conductors.

The flux-density in the teeth. This will depend upon the relative importance
of securing a high efficiency and of building a cheap machine. There is no doubt
that very high flux-densities can be employed without making the temperature

* Small oonverters up to 200 k.w., 600 volts are generally wound with wave-windings with
two cirouits in parallel. In rare cases it is desirable to use the Arnold singly re-entrant windins
described on page 512. In these cases the best way of finding the points on the winding to which
the slip-rings are connected is the method descriwd by Dr. S. P. Smith and R. S. H. Boulding
in the Joum. LE.E,, vol. 53, p. 232.

670

DYNAMO-ELECTRIC MACHINERY

Oitte..§S^Ar..i9/4. Tjpe o«i. .•VM MaTOir- ROTAmr ..../#. .

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300 KM A. leading at

fun ^0^4

ROTARY CONVERTERS 671

too high ; but if we consider the cost of power lost in the teeth we shall find in
most cases that it will pay, as an engineering proposition, to slightly increase the
size of the machine, so as to work at a lower density in the teeth and make a saving
in power. A density of 18,000 c.G.s. lines per sq. cm. is generally satisfactory
at 50 cycles. Where efficiency is specially important, a lower figure will be chosen,
and where efficiency is of less importance, a higher figure. Dividing 1-11 x 10® by
18,000, we get 6200 sq. cms. for the cross-section of all the teeth.

Number of slots. The fewer the numbers of conductors per slot on a rotary
converter and the greater the number of slots, the better from the commutation
point of view. From considerations of economy, however, we find it necessary to
group six or eight conductors in one slot on a 500-volt machine. Eight conduc-
tors per slot gives quite good commutation conditions where proper care is taken
in the design of the commutating pole. We therefore choose this number, and
arrive at 96-7-8 = 12 slots per pole.

Size of conductor. This will depend upon the power factor at which the converter
is intended to work. It is only by actual trial under the ventilating conditions
which obtain on any given machine that we can with certainty state the load
which can be carried by a conductor of a certain size at a certain power factor.
The considerations which determine the size of armature conductor on a converter
are given on page 544. Where the power factor \a lower than unity, the heating
of the conductors near the point where the taps are connected is very much greater
than the heating of intermediate conductors ; and the rate at which the heat is
conducted from the hot parts of the armature to the other parts is so uncertain
that no exact calculation is possible. It is found, however, that on 50-cycle con-
verters having a peripheral velocity of 30 metres per second, and with the means
of ventilation ordinarily available, one can, when the power factor is near unity,
work as high as 900 (nominal) amperes per sq. cm. ; that is, taking the current
as that of a o.c generator.

In actual practice, it is seldom found advisable to work the copper at such
a high current density as to bring up the temperature to the guaranteed tempera-
ture, because by so doing we should only be saving a small weight of copper at the
cost of considerable loss of power. The amount of material employed will depend
upon the efficiency which must be obtained. In the case under consideration,
if we have regard only to the mean temperature rise, we see from page 545 that at
a power factor of 0-96 the heating will be about 0-378 of what it would be on a
continuous-current generator. The current density, therefore, may be made

--.^ -—1-62 times as great. This would give us a possible current density of
vO"378

750 amperes per sq. cm. (nominal), and a total loss of 10,000 watts in the armature
resistance (see page 544). If we use a slightly greater section of copper so as to
work at 640 amperes per sq. cm. (nominal), we will reduce the copper losses by 1500
watts, and at the same time run less risk of exceeding the temperature guarantees
at the points of the winding near the taps. In checking the mean temperature
rise of the copper above the outside of the insulation, we have the following expres-
sions (see page 570), 0 000654x1 16x168x168x0 •378x8x1-4 = 90 watts per
metre length of coil. The cooling surface per metre length of coil is 840 sq. cms..

DYNAMO-ELECTRIC MACHINERY

r"" T T r T T

Tn. Bll.— ScMIoimI diawing of 12IiO K.v. S-pbuo rotary coi
Bi«dJlcatlan Ma. U, pato HO-

liiiiiiii' T'i'i'i'ni'ii

0 ' Itincka

ROTARY CONVERTERS

674 DYNAMO-ELECTRIC MACHINERY

giving 0-107 watt per sq. cm. As the thickness of the insulation is 0-13 cm. and
the conductivity 0-0012 (see page 225), we have

0-107 X 0-13 _.,, op
0 0012 ~ '

difierence of temperature between inside and outside of coil. As a matter of fact,
the temperature of the top conductor will be more than^^this, because it carries the
heaviest eddy current.

The air-gap. The air-gap under the main poles of modem rotary converters
is made quite short. It must not be so short as to cause excessive losses in the
pole faces due to the open slots. If it is made about half the width of a slot, and
if the pole \b built up of laminations, it will be quite long enough. A length of 0-5
cm. is enough for converters up to 150 cms. in diameter. For very large machines
the air-gap will be made a little greater for mechanical reasons. On 25-cycle con-
verters the pole pitch is usually greater and the ampere-turns per pole greater,
so that one usually has a rather bigger air-gap even up to 1 cm. for large machines.

Magnetization curve. The method of working out the number of ampere-turns
per pole for three different voltages, 500, 550, and 580, is shown on the calculation
sheet, page 570. In plotting these on a curve, it is best to take as ordinates the
flux-density in the gap rather than the voltage, so that the curve will be con-
veniently available for all machines built on the same carcass, whatever the voltage.

Shunt winding. The method of working out the shunt winding is exactly
similar to the method described on page 331. The ampere-turns at full load have
been taken at 6330 instead of 4230 to enable the poles to be over excited by 2100
ampere-turns. This is to make the rotary converter draw a leading current. The
effective armature-turns per pole when the converter is drawing full-load current
wattless are 7000 (see page 599), so to draw 0-3 of full-load current we will require
7000 X 0 3 = 2100. The cooling conditions on the shunt coil are worked out as shown
on page 231. The allowance 18-7 sq. cms. per watt is very liberal for a 50-cycle
converter, because as a rule there is a very great draught from the commutator
necks, which is diverted in a horizontal direction, and gives very good cooling
conditions.

Series winding. This is not strictly necessary on this machine, because a booster
is to be fitted for raising the voltage at full load. It will, however, be found that
the addition of a series winding greatly facilitates the work of the booster, and
enables a smaller machine to do the work. As the booster brings up the voltage,
the excitation of the converter ought to be automatically increased, so as to be
equal to the excitation corresponding to the voltage in question, as ascertained
from the magnetization curve of the machine. If there is no series winding, the
excitation will not be sufficiently increased at the higher voltages, and the power
factor will change towards the lagging side, so that the booster will have to be more
highly excited to bring up the voltage.

It is not necessary to have more than one turn on the series winding, and even
this will be shunted so that it will not carry the full-load current.

Commutating pole winding. The calculation of the commutating pole]^winding
will be understood from the formulae given on page 480, and from the dimensions

ROTARY CONVERTERS 675

of slots and pole given on the calculation sheet. In this case we have a strap coil
on the armature with a short throw.

J - _ 2-25 17-5 ^_
Ze = l-25x2^j:jX-3j-=0.57,

^ =

|IxO-46(log,o^~-0.2)=0-85,

Bc=2-8x 2-9 X '.''-^ = 2430,

4: -49

2>« + 66-Cft = 4-49.

As the commutating pole has not the full axial length, but only 15 cms., or 17*5
cms., allowing for fringing,

^^x 2430 = 4300= Be'.

17-5

If the length of air-gap under the pole is 1-37 cms., we have

0-796 X 4300 x 1 -37 = 4700 ampere-turns per pole ;

so that two turns, each carrying 2360 amperes, will be sufficient.

Commutator and bmsh gear. One of the difficulties in the past with 50-cycle
converters has been to get a great distance between the positive and negative
brush arms. The time taken for a commutator bar to pass over the pitch of the
brushes is only y^th second, so that if we make the pitch 25 cms. we get a cir-
cumferential speed of 25 metres per second (about 5000 feet per minute). This
speed is found to give satisfactory operation. If we make the measurement of the
brush holder on a circumferential direction rather small, we can get a clear 22 cms.
between brush holders, a distance quite great enough to cause the brush arms to
clear themselves if a flash-over should accidentally occur. A type of brush holder
which is good for this purpose is that illustrated in Fig. 515.

The diameter of the commutator in this case will be 112 cms. to give us 14 x 25
cms. of circumference. There will be 672 bars, giving us 11*5 mean volts per bar,
and 11 -5 ^0-72 = 16 volts actual. The length of the commutator depends upon
the number and size of brushes.

Width of brushes. From one point of view, when commutating poles are used,
there is an advantage in a wide brush, because it lengthens the time of commuta-
tion and lowers the voltage required to reverse the current. The brush, however,
must not be so wide that the arc moved through by the short-circuited coil extends
under the horns of the main poles. In order to make sure of the position of the coil
under commutation, with respect to the main poles and the commutating pole
at various stages of the motion under the brush, it is well to make a paper model
of the bars and coils and rotate them on a drawing of the brushes and poles. If
this is done with the machine under consideration, it will be seen that it is not
wise to make the brushes much wider than 2 cms., or the short-circuited coil will not
be sufficiently under the control of the inter-pole, but will sometimes be moving in
a field of the wrong value. This circumstance limits the width of the brush.

DYNAMO-ELECTRIC MACHINERY

ROTARY CONVERTERS

677

12 Inches

O 10 20

' ■■' - '

M I'M 'I I II 'I I

SO cm

2/V

Fig. 510. — Longitudinal section of rotary converter and A.O. booster, mounted on the
shaft between the converter armature and the sUp-ringa.

W.M.

678 DYNAMO-ELECTRIC MACHINERY

Brush contact area. While it is recognized that some kinds of carbon brushes
work well up to densities as high as 10 amperes per sq. cm. (65 amperes per sq. in.),
experience shows that machines with ample brush capacity give the least trouble.
There appears to be very little to be gained in making the density less than
6 amperes per sq. cm., and that may well be taken as a standard for traction
rotaries where we do not wish to cut down the cost to the smallest possible amount.

If we put six brushes per arm, each with a contact area 2x4-5 cms., to collect
the 336 amperes per brush arm, we get a density of 6-2 amperes per sq. cm. —
quite a suitable figure. At 50 per cent, over load we will have a little over 9 amperes
per sq. cm.

Grade of brush. For a high-speed commutator it is desirable to choose a brush
with a great deal of graphite in its constitution. The brush, however, should not
be so friable as to wear badly in the holder. It should be fitted with flexible pig-
tails of low resistance, capable of taking heavy over-loads without overheating.
Many of the graphitic brushes of sufficiently solid composition give a fairly low
contact voltage drop, and we may allow 1-7 volts for the drop in positive and
negative brushes taken together. If we use a good metal-carbon brush on the slip-
rings, taking care to have no chattering, and to keep the brush well in contact
with the ring, we can get the contact drop per pair of rings as low as 0-7 volt, equiva-
lent to 1 volt c.c. This enables us to estimate the total brush loss by multiplving
the continuous current by 2-7. In practice, however, it will often be found that the
total brush drop over positive and negative brushes is as high as 2-5 volts, and if the
slip-rings are not working well we may have an additional volt lost there.

Efficiency. According to the specification, the machine will be judged by its
calculated efficiency and not from its efficiency measured from input and output.
We may take the losses as given on the calculation sheet. The figure 21 k.w.
for friction and windage includes 4 K.w. for brush friction. The iron loss with
ordinary good dynamo sheet steel ought not to exceed the values calculated on
the sheet, because Fig. 29 gives the losses rather on the high side. In taking the
field loss, we should take the shunt excitation rather less than 6330 ampere-turns
per pole, because the series turns contribute a substantial amount. We may take
the shunt excitation at 4000 for 625 volts. That will give us 6-15 amperes at 525
volts, or say 3-3 K.w. loss in shunt and rheostat. We multiply the armature
resistance by 0-378, and to this add the resistance of the commutating poles and
the diverted series winding. This combined resistance multiplied by 2360* gives
us 11 -7 K.w. loss at full load, and the other figures set out on the calculation sheet
at the other loads. The brush losses may be conveniently obtained by multiplying
the armature current by 2-7.

Design of the amortisseur. It appears from the Specification No. 14 that the
A.c. power is supplied by steam turbines. As the speed of the steam turbine is
very uniform, we will not be troubled with an unsteady frequency, so that the
dampers may be of the very simple character shown in Fig. 511. They are built
up of three round rods passing through holes near the pole face, and two shaped
bars flanking the poles, the whole connected by two stout copper bars. In this
case the dampers are not connected from pole to pole, as this is not necessary. If
the source of the a.c. supply had been such as to give us an imsteady frequency,

ROTARY CONVERTERS 679

more elaborate precautions would be taken to obviate hunting. The matter is
considered further on page 601.

Direct-connected exciter. This converter is intended to run at certain times
from the c.c. side as a motor and supply alternating current from the transformer
to the high-tension mains. If it ia intended that it shall always be in parallel with
a synchronous generator, when it is running in this way, no exciter for the converter
will be necessary. But if it is intended that a converter or converters shall be the
only machines supplying the power to the a.c. system, then it will be necessary to
excite each converter by means of an exciter driven by itself for the purpose of
keeping the frequency nearly constant. If no such exciter or equivalent device
were employed, there would be nothing to maintain the speed of the converter.
A heavy lagging load upon the a.c. side would weaken the field-magnet of the
converter, and the speed would rise. With rising speed the inductive load would
call for more lagging current, and the converter might run away. If a direct-
driven exciter, with an unsaturated field-magnet, is used to excite the converter,
a slight change in the speed of the set raises the exciting voltage by a large per-
centage, and so corrects the tendency for the speed to go up. One should build
the exciter with the saturation of the field-magnet at its working voltage well below
the knee of the magnetization curve. Such a machine, when increased in speed
by 1 per cent., will give a rise in voltage of 5 per cent, or more.

THE DESIGN OF AN AC. BOOSTER.

The driving of the booster. An a.c. booster for changing the electromotive force
supplied to the taps on a rotary converter should be mounted on the shaft of the
rotary, so as to ensure perfect synchronism. So long as the output of the booster
is only 10 or 15 per cent, of the output of the rotary, the power required to drive
it when-raising the voltage can be supplied by the converter, which then runs partly
as a synchronous motor. When lowering the voltage the booster runs as a motor,
and the converter then acts partly as a c.c. generator. It is found in practice
that the commutation of a suitably designed converter is not interfered with for
small ranges of boosting, though in theory there is an unbalanced armature reaction
in the armature, in so far as it acts as a motor or generator. When the booster
is raising the voltage the converter will be acting as a motor to a certain extent,
and this will give an armature reaction which will assist the commutation. When
the booster is lowering the voltage the rotary acts as a generator, and the armature
reaction opposes the commutation, but the resistance of the brushes is sufficient
to prevent either of these effects becoming apparent where the output of the booster
is small as compared with that of the rotary. This matter is one, however, which
must be borne in mind where the percentage of voltage variation required is very
great. It is, of course, possible to drive the booster by an independent synchronous
motor. Where this is done great care must be taken to secure true synchronous
running, or the effect on the commutation and regulation of the rotary will be
disastrous. A synchronous motor for such a purpose should be designed with a
very strong field, and provided with a very heavy damper. Such an arrangement
is not to be recommended unless the frequency of supply is very steady, because

580 DYNAMO-ELECTRIC MACHINERY

any phase-swinging which might occur on the rotary and booster would at times
get out of step, and cause bad fluctuations in the rotary voltage.

The usual practice is to drive the booster by mounting it on the shaft of the
converter, either between the slip-rings and the converter's armature (in which
case one must have a rotating-armature booster), or outside the slip-rings (in which
case it is best to have a rotating field). The advantage of the first arrangement is
that it makes a very compact machine with no extra terminals, except the field-
terminals of the booster. The outer ends of the booster armature winding are
connected directly to the slip-rings, and the inner ends directly to the taps on the
rotary armature (see Fig. 505). The field-frame of the booster can be conveniently
mounted on the yoke of the converter (see Fig. 516). This arrangement is to be
recommended in all cases where the booster is required to be continually in circuit.
There are, however, some cases where the booster is only required occasionally,
and where it is desirable to cut it out of circuit at times when it is not wanted.
In these cases it is more convenient to mount the booster outside the bearings of
the converter. The rotating field of the booster can generally be over-hung. If
the booster is wound for six phases, six terminals would be required to take the
current into the stationary armature, and six more to take the current out. It
is therefore desirable to arrange for a stationary armature booster to be connected
in three phases only. This will be possible on a six-phase rotary, if we do not
inter-connect the middle points of the transformers so as to make a six-phase star.

Size of frame. The case of the over-hung booster will, in practice, be found to
be the exceptional case. Most boosters will be built with rotating armatures placed
between the slip-rings and the converter armature. We will choose this type of
machine for the booster, which we will design to meet Specification No. 14. It
will be found convenient, and, on the whole, more economical, to develop a frame
of a certain diameter to be used as a booster in connection with a certain diameter
of rotary. A common output for a booster is 10 per cent, of the output of the
converter, so we should choose a diameter of the booster frame which will make an
economical machine for that output. Smaller outputs will then be put on the same
diameter with the frame shortened, and larger outputs on the same diameter with
the frame lengthened. This plan, though calling for more material than would
be necessary with the theoretically best diameter, will be foimd to work out, on the
whole, most economically when the development expenses are taken into account.
Moreover, it is not worth while to cut down the material to the smallest possible
amount, because in any case the cost of the booster, though high in comparison
with its output, is a small percentage of the total cost, and a little more or less
material hardly affects it ; whereas if we put in a good large section of copper in the
armature we reap the benefit in the increased efficiency. If we were to cut down the
armature copper to the smallest amount that would meet the temperature guarantees,
we would make the PR losses in the booster nearly equal to the PR losses in the
rotary itself.

In the case under consideration, the machine is required to bring down the
voltage to 460 and to take it up to 550. For traction work it is to compound from
525 to 550 volts. In the compounding it will be assisted by the series winding, so
that if we aim at 45 volts c.c, or 28 volts three-phase, it will be sufficient. That

ROTARY CONVERTERS 681

is to say, the output of the booster will be about 9 per cent, of the output of
the converter.

For a six-phase converter with a revolving armature booster we will usually
have a six-phase armature (if placed between the slip^rings and converter arma-
ture), there being as many circuits through the booster as there are taps on the
converter. The scheme of connections is that shown in Fig. 505. For a multipolar
machine there will be as many paths in parallel in each of the six phases as there
are pairs of poles. Thus, in the case under consideration, there will be 6 x 7 = 42
paths through the booster. This makes the determination of the constant Ke a
little perplexing, unless we adopt a rule which takes us from the six-phase case to the
three-phase case. We may argue as follows : Consider how many conductors would
be required on an ordinary three-phase, star-connected armature, and take this
number as the Za in the formula (1) given on page 24. These conductors will form
three of the legs under two poles in Fig. 505. We will ultimately have to find room
on the armature for the other three legs, but they will not add to the generated
E.M.F. If we have more than one pair of poles, we will have as many paths in
parallel per phase as there are pairs of poles.

In the example under consideration we want 45 volts o.c. or 28 volts three-
phase. Now, the field being stationary, we can, in general, have a rather wider
pole face than we would have on a revolving field, so we will take the constant
Ke at 0-41 (see page 30).

The calculation sheet is given on page 582 and a drawing of the booster on
Figs. 512, 514 and 516.

As the method of working through the calculation sheet is so similar to the
method described in connection with the a.c. generators described on pages 321
and 348, it is not necessary to go through it in detail. We will just refer to those
points which are special to this machine.

We have chosen 168 slots or 12 per pole, and we have 2 conductors per slot.
This would give us 48 conductors per pair of poles, or 8 conductors per phase for
a six-phase machine. It is, however, convenient to have an odd number of con-
ductors per pole, because we wish one terminal of a coil to be on the outer end of
the armature for connection to a slip-ring, and the other terminal to be on the
inner end for connection to the converter armature. We therefore choose 7 con-
ductors per phase, and leave out one of the 8, a piece of treated wood taking the
place of the 8th conductor in the slots. We thus have 7 conductors in each branch
of the star ; that is to say, 21 conductors on a three-phase machine. Our voltage

formula then becomes

28 = 0-41 X 715 X 21 x AgB x lO"®,

^f,B=0-455xl08.

Although we only count 21 conductors for the purpose of this formula, we must
provide room for another 21, which form, as it were, another three-phase machine
with the phases displaced by 180 degrees from the first set of conductors.

It will be seen that on the size of firame taken the magnetic loading and the
current loading are both quite light. With only 17,700 c.G.S. lines in the teeth,
and the current loading only 170 amperes per cm. of periphery, it is not necessary
to work out the cooling conditions in the armature.

582

DYNAMO-ELECTRIC MACHINERY

KV.A..MS. ; P.F ; Phwe 6. ; Vo\tM.2B.:.»^n^.i Amps per ter

H.P. Amps p. cood. J.G.iff ..AmpB p. br. aim Temp, riae

//eo.

...iffT. .Poles* Emc. Spec. ■*••«%...

Cycles .v^iC^:..; R.P M.^i?^...; Rotor Amps
.Regulatioa Overiood2)>/ff.

easterner.

Order No ; Qnof' No. ; Perf. Spec ; Fly-wheel effect

Frame 92^

Air

CocoxilZoS ', Gap Area

28 voits=:4yx 7:/5

<'7<5^C.*V'55...

; poaa. laZ

Zi .:^S5

4^9,QOO jCircnm.

J7Q

MJLxRjPli
K.V.A.

^5-7* 10^

Arm. A.T. p. pole.

BCaac Fid. AT.

Armatupe. Rev. Qtat

Dia. Outs.
Dia. Ins

Gross Length
Air Vents — -

Opening Min Mean

Air Velocity

Net I,ength iS x-Sg

Depth b. Slots

Section Z/B -.Vol

Flux Density ^

Loss:/2fi.p. cu.C/ZZ. Total

Buried Cu Total

Gap Area ; Wts

Vent Area ; Wts

Outs, .^rea : Wts

No of 'Segs
No of Slots
K. -

/JMn.

^2l

JS4u

iiaon

V29na

Circ.

6S^

Section Teeth .
Volume Teeth.
Flux Density.

fgsoo

Loss'/J^Sp. cu CflL-Total

/7.700

■o
c
o
o

Weight of Iron.

Star or Mesh Throw

Cond. p Slot

Total Conds

Size of Cond. :*3_x
Amp. p. sq.jC^.

Length in Slots /^

Length outside j2i2-Sum
Total Length

Wt of x.ooojS^i^Total
Res. p. 1.000 :#■ -Total

Watts p

Surface p —

Watts p. Sq

270

62
7331

2^0_

16QI1.

.-^

2S4:^

3Q7

J±L

^±.

/4r Poles

ffr

)s ^

?'^

/'72-
1

■

Field Statjdr Rotftf.

Dia. Bore _.,^_^_.

i Total Air Gap

Gap Co-eff. K, ^—

Pole Pitch2C:7Pole Arc
Kf

92-6

muxi>erVole2'38^/0^

Leakagen.1 f.1 *S

PiXf^ZOS Flux density

Unbalanced PuU

lA.

i'(2

H-'S

286 '^/qE.

MQOn

No. of Seg. .
No.ofSlotsL

Vents L

Mn.Cizc.

.Section

Weight of Iron.

f-2

/6S Slots

'S^HZx3750X'7S6^2340

Shunt.

A.T.pPolen.Load

AT. p. Polef.LoadI 3384- 2600 '

Surface .%000 (B,6Q0

Surface p. Watt l6 /g .

LR.

Amps.

No. of Turns.
Mean 1. Turn .
Total Length.
Resistance

^,W '65

A:2S- 93S*2

SCO

70 7</

\290Q 29m

Res, per 1. 000,

Size of Cond

Conds. per Slot.
Total

l^U2^ 'OS
015 sqcm 4^. cms.

Length

Wt. per i.ooo-
Total Wt

IQBhotjnOCOiSL

Watts per Sq

Star or Mcah

Paths in parallel

/3 ^ i Jjggg,

tOS I 10^

Magnetization Curve.'

Core

Stator Teeth
Rotor Teeth
Gap

Pole Body
Yoke

Section.

Z53a

52QA
2QS.

2^0 20

L«ngth

:iL

.-25..voit3.

iLLp-^w

A.T.

\J126. iTfo^ SO
\2Q9Q.87Sd
fZSOi> 7\'5U2Q. CS,<M 9

\/a70p /.^-L2g^ l2jmJ3^
^2596

ernciENCY.

Friction and W.

Iron Loss

Field Loas

Arm. &c. I*R— .
Brush Loss _

Ijloid.

I:A.

4'S

Fall.

±Ji

±A.

rs_

ZS I '6

/O'^

Output
Input

Efficiency ^L

/i±

12^

.J2d. Volts.

A.T.P <•* A.T.

^8aa

300.

_i_ _i L_

t'S rs rs

4 5 I ^-5 : 45

i'9

:s.

±S.

i'9 ' f'9

1^ ±

^164-

. J^.Voits.

\23Mm3m

^44-

i9,00i' tSO

Ll.pem

f^ooc 21

i270C /6'S

A.T.

_M2

33g

.S3g.

3926

Commutator.

Dia.

Ban _
Volts p. Bar.
Brs. p. Aim _
Size of Bis. .
Amps p. sq. .
Brush Loss .
Watts p. Sq. .

.Speed

Mag. Cur.

Perm. Stat. Slot
,. Rot. Slot X
„ Zig-zag
X

Loss Cur.

2 X

177
End

X
X X

Amps ; Tot.

;X. -

; r- «=

+ —

Imp. V +
Sh.cir. Cur

Starting Torqu?
kax. Torque _

Max. H.P

Slip

Power Factor

ROTARY CONVERTERS 583

Shunt winding. Some difficulty is sometimes experienced on these booster
generators in finding room for both series and shunt turns. As can be seen from
the figures for the magnetization curve, the shunt ampere-turns to give 28 volts
at no load are 3164. At full load the teeth will require about 220 ampere-turns
more, so that we have taken 3384 as the ampere-turns to be provided by the shunt
at full load. Under the specification the load must be slightly leading, so that the
armature reaction will assist rather than oppose the shunt ampere-turns.

We have allowed 16 sq. cms. per watt on the shunt coils. This is sufficient, in
view of the very good ventilation induced by the armature of the converter. We
find that 800 turns per pole, of wire having an area of 0-015 sq. cm., will be required.
Thus the exciting current at full voltage will be 4-25 amperes.

Series winding. We find that we require at full load about 2800 ampere-turns
per pole for the series winding ; we must therefore have more than one turn. We
might put two turns per pole and divert a great part of the 2360 amperes. A rather
nicer arrangement is to put three turns per pole and put two paths in parallel.
We can then send one half of the current one way around the frame and the other
half the other way around the frame, and thus avoid magnetizing the shaft, as we
would do if we passed a large current once around the frame.

The figures for the efficiency will be found on the calculation sheet.

LARGE LOW- VOLTAGE CONVERTERS.

We will now give a specification for a 2000 k.w. rotary converter intended for
electrolytic work.

684

DYNAMO-ELECTRIC MACHINERY

SPECIFICATION No. 15.

Extent of
Work.

Oeneral
Purposes of
riant.

2000 K.W. ROTARY CONVERTER, 250 VOLTS, 50 CYCLES.

240. This specification provides for the manufacture,
delivery on site, erection, testing, and starting to work of
rotary converters and transformers, together with all acces-
sories and details as hereinafter specified, in the sub-station
of the Company hereinafter called the Purchaser, at

241 . The duty of the plant will be to convert 3-phase power
supplied at a voltage of 11,000 at a frequency of 50 cycles
per second into continuous-current power at from 230 to 270
volts, to be used in electrolytic work. The plant must be
suitable in all respects for this purpose.

Characteristics 242. Ihc rotarv COUVC

of Rotarj'. - , • , • ^

characteristics :

rter shall have the

Normal output when

running a.c. to c.c.

2000 K.v.

Number of phases

Frequency

50 cycles per second.

Normal continuous-

current voltage

250

Continuous-current

amperes

8000.

Kind of excitation

Shunt wound.

Adjustment of voltage

on rheostat

230 to 250 Tap (1).

240 to 260 Tap (2).

250 to 270 Tap (3).

Leading idle power re-
quired when running
A.c. to c.c, at the
highest voltage on
any tapping

Over-load capacity

600 K.v.A. leading wattless.

25 per cent, for 3 hours at unity
power factor. 50 per cent,
for 10 minutes at imity power
factor.

ROTARY CONVERTERS 685

Temperature rise after
continuous full-load
runs at 250 volts,
0-955 power factor
on H.T. side 40° C. by thermometer.

Temperature rise after
3 hours 25 per cent,
over load 55° C.

Puncture test 23,000 volts alternating at 50

cycles applied for 1 minute
between transformer high-
tension windings and frame.

1500 volts alternating at 50
cycles applied for 1 minute
between all low-tension wind-
ings and frame.
Mean voltage between
commutator bars not
to exceed 11 volts.

243. The rotary converters are only intended to run from a.o. to o.o.
A.c. to c.c.

244. The 6-phase rotary converters shall be of the two- Type,
bearing horizontal type with shunt-woimd field magnets and
commutating poles.

245. The speed shall not exceed 250 r.p.m. speed.

246. They shall be fitted with commutating poles. commutating

247. Each rotary converter shall be connected by cables comiectionB.
running directly from the slip-ring brushes to the l.t. trans-
former terminals without the intervention of any switch gear.

248. It shall be started by means of a starting motor starting.
direct-connected to the shaft, and shall be synchronized on

the high-tension side of the transformer.

249. The efficiency of each rotary converter shall be Efficiency.
calculated from the separate losses, which shall be measured

in the following way :

(a) Iron loss, friction and mndage. The machine shall
be run as a c.c. motor at full speed and at various

586 DYNAMO-ELECTRIC MACHINERY

voltages, and measurements shall be made of the c.c.
power taken to drive it, and of the exciting current taken
at various voltages.

(6) Copper losses. The resistance of the armature of
the rotary and booster and of their field windings shall be
taken by measuring the voltage drop in them when a
substantial current is passed through them. From these
resistances, after making due allowance for the observed
temperature rise, the PR losses shall be calculated on
the assumption that on a 6-phase converter the armature
copper loss is 0'3 of the copper loss when loaded as a
CO. generator.

(c) Brush losses. The brush PR losses shall be taken
as equal to the number of watts obtained by multiplying
the continuous current delivered by the converter by
3 volts, unless the Contractor can demonstrate to the
satisfaction of the Purchaser that on his machine the
sum of the brush losses on the c.o. side and the A.c. side
is substantially less than this, in which case the actual
voltage drops on the commutation brush and slip ring
brushes shall be taken.

{d) Transformer iron losses. These shall be taken by
measuring the number of watts suppUed to operate the
transformer at full voltage, 50 cycles at no load.

(e) Transformer copper losses. These shall be taken
to be equal to the power required to circulate full-load
current through the transformer when short circuited.

G?MaStee ^^^- "^^^ Contractor shall state what calculated efficiency

he is prepared to guarantee on the above basis. He shall
guarantee that the combined plant in commercial service
shall have an overall efficiency not more than 1 per cent,
lower than the calculated figure.

Houra^?' 251. The converters are intended to nm in almost con-

annum, tinuous service, and each machine will probably run at

approximately full load for 7000 hours per annum.

Value of 1 per 252. The cost of elcctrical energy may be taken at approxi-
EfflcilS?y. " mately 0-25 pence per kilowatt hour. So the value of each
ifper cent, in efficiency is about £145 per annum.

Bonus and 253. lu view of the great importance of high efficiency

the Purchaser will pay a bonus of £50 for each ^^th per cent

ROTARY CONVERTERS 687

by which the eflSciency of the combined plant (transformer and
converter) shall exceed the guaranteed calculated efficiency,
and the Contractor shall pay a penalty of £50 for each y^^th
per cent, by which the efficiency shall fall below the guaranteed
calculated efficiency.

If the efficiency of the plant shall fall below 92 per cent.,
the Purchaser shall be at liberty to reject it.

Figures for the calculated efficiency shall be given for
full load, three-quarter load and half load, both at unity
power factor and at 0-96 leading power factor measured
on the H.T. side, but the bonus and penalty shall only be
paid in respect of the efficiency at full-load, unity power
factor.

254. Tappings shall be provided on the high-tension side ^^^^^JreT
of the transformers, wherebv the range * of voltage obtainable Means of
on the c.c. side shall be from 230 to 250, or 240 to 260, or Tappings.
250 to 270. Under these conditions the rating of the plant
when running on the 230 to 250 tappings shall be taken to
be 1900 K.W.

255. On the middle set of tappings the voltage shall be
varied from 240 to 260 by changing the excitation of the
converter, and rheostats shall be provided to enable the range
specified on each set of tappings to be obtained whether
the machine is hot or cold.

266. The plant shall be so designed that by means of variation of
regulating field rheostats the power factor in the h.t. side
at full load can be varied from unity to 0-95 leading, the
pressure on the h.t. supply being 11,000 volts and the periodi-
city being 50 cycles per second. When the power factor
is 0*95 leading, the Voltage may be above the middle voltage
obtained from the tapping, but it must not be higher than
the highest voltage in the range specified.

* If the converter were to be provided with a booster, the following claiue would
be inserted instead of 254, 255, and 256 :

256a. The plant shall be so designed that when running under working conditions, Variation of
supplying current for electroljrtic purposes, the continuous-current pressure shall Voltage,
remain steady so long as the high-tension a.c. pressure shall remain steady, and the
adjustments remain undisturbed. It shall, however, be possible to obtain on the c.c.
side any voltage from 228 to 275 by the adjustment of the booster and rotary converter
rheostats, so long as the high-tension pressure is maintained at 11,000 volts, and the
frequency at 50 cycles. It. shaU, moreover, be possible to maintain the power factor
of 0'95 leading with any c.c. voltage between 240 and 275 volts.

588 DYNAMO-ELECTRIC MACHINERY

otL^ ov« 267 . It shall be possible by means of the rheostats provided
to change the load from one rotary converter to another
without alteration of the c.c. voltage of supply.

gwd^ 258. Two field regulating rheostats of approved type shall

Eheoetat. bc providcd for each rotary converter. One of the rheostats
for the rotary converters shall be placed adjacent to the
high-tension panel for convenience of adjustment while the
converter is being synchronized, and the other shall be placed
adjacent to the c.c. panel for convenience of adjustment
steps. of the c.c. voltage. The steps in this latter rheostat shall

be so fine that it shall be possible to obtain over the whole
range variations not exceeding 1 volt per step. The rheostats
shall be supplied with the necessary face-plates, bevelled
gearing (if necessary), and hand- wheels.

starting Motor. 259. Each Starting motor shall be of the squirrel-cage
induction type mounted on an extension of the armature
shaft and suitable for running on a 3-phase, 50 cycle, 440
volt circuit. This circuit will be provided by the Purchaser,
and fed from independent transformers supplied under another
contract. Each starting motor shall be of ample capacity,
and shall be capable of driving the rotary converter at full
speed, normal excitation, for at least 20 minutes, so as to
enable the high-tension side of the static transformers to be
synchronized with the high-tension system, under all con-
ditions of commercial operation. They shall be so designed
that when a converter is running at synchronous speed the
A.c. voltage generated by it shall not differ from the normal
supply voltage by more than 15 per cent, either way, and
provision shall be made so that the slip of the starting motors
can be readily varied to meet this condition.

Absence of 260. Uudcr auy of the conditions of voltage and periodicity

Hunting. contemplated in this specification, and under practical
working conditions, the rotary converters and transformers
shall run parallel with one another, with the rotary con-
verters now existing in the sub-station, and with the high-
tension system, without hunting or falling out of step.
This steady operation shall be maintained notwithstanding
the fact that the load may be fluctuating from no load to
50 per cent, over load, and that the total load is unequally
divided between different machines in the sub-station.

ROTARY CONVERTERS 589

261. Each rotary converter shall operate sparklessly under commutation.
all normal working conditions from no load to 50 per cent.

over load. It shall be possible to obtain 100 per cent, over load
for 2 minutes without causing such sparking or heating as
will injure the brushes or commutator. The above conditions
as to operation shall be met without rocking the brushes.

262. The commutator of each rotary converter shall be commutators.
designed with a view to low-temperature rise and perfect
coromutation. It shall consist of hard-drawn copper seg-
ments insulated from each other and from the frame by means

of mica. The mica between the segments shall be of such
quaUty that it wears at the same rate as the copper. The
wearing depth of the commutator on the machine put forward
shall be stated ; and due preference will be given, when
considering tenders, to machines having great wearing depth.
The proposed peripheral speed of the coromutator shall be
stated. No peripheral speed of the commutator is here
specified, but no tender will be considered unless the tenderer
can show satisfactory results obtained on similar machines
with as high peripheral speed as proposed in his tender.

263. Each tender shall show by means of drawings the o.c. and a.o.
type of c.c. and a.c. brush gear proposed and the arrange- ^'
ments for supporting it. All brush gear must possess the
following characteristics : (a) The supports must be very-
rigid and not Hable to alteration of position ; (b) Each brush

must slide in a manner truly parallel to a given direction ;
(c) It shaU sUde or move without more friction than is neces-
sary to prevent chattering, and the amount of friction must
be reasonably uniform ; (d) The current shall be led out of
or into each brush by means of flexible connections ; (e) The
positions of the brush holders on the brush arms shall be
staggered so as to prevent uneven wear on the length of the
commutator or sKp ring ; (J) The holders shall be designed so
that adjustments of position and pressure are easily made
when the machine is running, and so that the brushes can be
easily removed and replaced ; (g) The parts shall be so shaped
that they shall not be injured by an arc if the machine
flashes over.

264. Each rotary converter shall be provided with an osciuator.
apparatus for keeping the armature moving backwards and
forwards axially, so as to prevent the formation of ruts on

the commutator and sUp rings.

590

DYNAMO-ELECTRIC MACHINERY

Type of
Brushes and
Flexible
Connections.

265. The type of brushes to be used on the commutator
and on the slip ring, and also the type of flexible connections,
shall be stated. Tenders shall also state the shortest distance
from one o.c. brush arm to another.

Current
Density in
Brushes.

266. The current density in the brushes at full load shall
be stated, and in considering tenders, preference will be given
to designs having a smaU current density.

Terminals and, 267. All counectiug straps which bring the current from.
ons. ^^^ brush holders to the terminals must be bolted together so
as to break joint and present large conducting surfaces where
the current passes from one strap to another. The terminals
shall be designed to fit into the copper straps provided by
the Purchaser for conducting the current from the rotary to
the switchboard, particulars of which will be supplied. The
design of these terminals must meet with the approval of
the Purchaser. Independent terminals shall be provided for
the ends of the field windings. The connections between the
field coils shall be very substantial, there being no loose
unanchored wires free to vibrate.

Bearings.

268. The bearings shall be of the self-lubricating type,
preferably with revolving oil-rings. They shall be self-
aUgning and spHt horizontally. The bottom bearing shaU
be arranged so that it can be removed without raising the
shaft more than 0*1 inch. It must be possible to remove
either bearing cap without dismanthng the brush gear.
The bearing pedestal shall be provided with oil gauges and
drain cocks. The oil wells shall be so covered that no dust
can enter, and the caps of the oil wells must be made so that
they are not detached from the housing when opened. The
design of the journal oil throwers and oil catchers must be
so perfect that no oil is visible outside the bearing after a
six hours' run.

Insulation.

269. The insulation of all conductors lying in slots must
consist largely of mica, and must be treated to prevent
deterioration due to moisture.

Sample Coil.

270. A sample armature coil similar to that proposed to
be used on the machine in question, showing the arrangement
of straps and insulation, though not necessarily of the same
size, must be submitted with the tender.

ROTARY CONVERTERS 591

271. Arrangements shall be made so that commutator Grinding Gear,
grinding gear of approved type can be readily j&xed to the

rotary converters.

272. All foundations, ducts, trenches and covers will be Foundations.
provided by the Purchaser. Within four weeks of the
placing of the order for the machinery the Contractor shall
supply to the Purchaser sufficient drawings and templates to
enable the Purchaser to lay out the foundations. If sufficient
information is not so supphed, any alterations or additions

to the work on the foundations shall be done by the Con-
tractor or by the Purchaser at the Contractor's cost. The
Contractor shall be responsible for the levelling and grouting
in of his machinery on the foimdations provided.

273. The Contractor may use at his own risk for the erection ^^ of crane,
of the machinery the overhead travelhng crane provided by

the Purchaser when the same is not required for other purposes.

(See Clauses 8, p. 271 ; 55-59, p. 379 ; 125, p. 461.) Accessibility.

274. All connections to existing bus-bars, machinery, etc., Tjmejfor
A\ be ca
Purchaser.

shall be carried out at such times as are convenient to the clmn^ions.

275. As each part of the machinery is erected, it shall be checking of

• Work

passed by the engineer of the Purchaser or his authorized
representative ; but such passing shall in no way exonerate
the Contractor from any responsibiUty under his guarantee.

276. All the working parts of the apparatus supphed shall interchange-
be made to gauge so that corresponding parts shall be inter-
changeable wherever possible.

277. All screw threads shall be of Whitworth's standard, screw Threads.

278. Before despatch from the manufacturer's works, the Painting.
rotary converters shall have all rough places filled, and shall

be painted with one coat of the best paint. After erection
on site they shall be given two final coats of paint of an
improved colour, and finally varnished in the best manner.

279. Cables or other leads connecting the transformer cables, etc.
L.c. terminals to the shp rings will be provided by the Pur-
chaser. The terminals on the shp-ring brush-holder supports

shall be provided by the Contractor ; they shall be very
substantial and of an approved shape.

592 DYNAMO-ELECTRIC MACHINERY

Spare Parts. 280. The tender shall state what spare parts are recom-

mended ; a Hst of such spare parts with their prices shall be
set out in a schedule.

^^i^l^Q^ 281. The tender shall contain a schedule giving the dates

at which the various sets covered by this specification can
be delivered and put into commercial service.

i^toS^' ^^2- "^^ following is a Hst of drawings attached to this

specification :

Proposed general arrangement of sub-station, drawing

No. .

Diagram of main connections, drawing No.
s^^i^ ^^^ ^^^' ^^® following is a list of drawings, etc., required with

required. thc tcudcr :

1. Outline drawings showing the apparatus to be
suppHed in plan and in elevation.

2. Drawings showing arrangement of c.c. brush-holder
arms and commutator, and showing construction of c.c.
and A.c. brush gear.

3. Drawings showing foundations with the position and
size of foundation bolts. These drawings should also
show where the a.c. and c.c. terminals of the rotary
converter will be placed.

4. Sample armature coil.

Testa. 284. The following tests will be carried out on the rotary

converter set in the presence of the Purchaser's representative :

1. Iron loss, friction and windage measurement, as
detailed in Clause 249a.

2. Measurement of resistances of rotary converter
armature and field, as detailed in Clause 2496.

3. Measurement of transformer iron loss as detailed
in Clause 249(Z.

4. Measurement of transformer copper loss as detailed
in Clause 249e.

5. Puncture test of 23,000 volts, alternating at 50
cycles, shall be appUed between all high-tension conductors
and earth for one minute.

Puncture test of 1000 volts, alternating at 50 cycles,
shall be applied between all low-tension conductors and
earth for one minute.

ROTARY CONVERTERS 693

During these tests all windings and connections other
than those to which the test is being applied shall be
connected to the core.

6. Temperature test. After erection the rotary con-
verter sets shall be run at full load in ordinary commercial
service until the temperature of all parts has become
substantially constant. The temperatures shall then be
taken of the armature copper, armature iron, commutator,
brush gear, field coils, and sHp rings, by means of thermo-
meters. The temperature of the air shall at the same
time be taken within three feet of the machine in line
with the shaft and at both ends of the shaft ; and the mean
of these two readings shall be taken as the temperature
of the air.

The temperature rise of the transformer oil shall be
taken as the difference between the temperature in the
hottest place available, and the temperature of the air
in the transformer chamber taken liiree feet from the
ground.

7. Efficiency. If the efficiency calculated from the
foregoing tests shaU be within the guaranteed figures, the
fact will be taken as prima fade evidence that the efficiency
guarantees have been met. If, however, the Purchaser
can conclusively prove, by means of properly calibrated
wattmeters connected in circuit after the plant is installed,
that the efficiency is 1 per cent, lower than the guaran-
teed figures, after making due allowance for losses in
cables and connections, it will be incumbent upon the
Contractor to amend the plant so as to obtain an over-all
efficiency within 1 per cent, of the guaranteed figures ;
and if it shall be impossible to so amend the plant, he shall
pay a penalty to the Purchaser of not more than £ for
every 1 per cent, below the guaranteed figure. Provided
always that if the efficiency shall fall below 92 per cent,
the Purchaser shall be entitled to reject the plant.

285. The Contractor shaU give to the Purchaser or his Notice oi
representative seven days' notice of any tests which he pro-
poses to carry out in the presence of the Purchaser or his
representative. Two copies of the results of all such tests
shall be furnished to the Purchaser.

W.M.

694 DYNAMO-ELECTRIC MACHINERY

DESIGN OF A 2000 K.W. ROTARY CONVERTER FOR ELECTROLYTIC WORK.

As the current per terminal on the c.c. side is 8000 amperes, one of the main
considerations in settling the design of this machine is the fixing of the number of
poles. The greater the number of poles, the fewer the amperes to be commutated
at each brush-arm, and the slower the speed. If first cost and efl&ciency were of
no importance, we would make a large number of poles. If we chose 32 poles
we would have only 500 amperes per brush-arm, which, at a voltage of 250, would
be quite easy to commutate. The length of commutator would then be some
16 inches, igid the cooling conditions would be good. However, 32 poles would
give us a speed of only 187 b.p.m. ; the size and cost of the machine would be rather
greater than really necessary. Experience shows that on the machines which are
to carry a steady load at a voltage of about 250, it is possible to go to 800 amperes
or more per brush-arm without endangering the commutation. It would, therefore,
be quite possible to choose only 20 poles and yet make a very good machine, but the
specification in this case says that the speed shall not be higher than 250 B.P.M.
We must, therefore, have at least 24 poles in this case. We will have 666 amperes
per brush-arm at normal load, and 835 amperes per brush-arm at 25 per cent,
over load.

As it is important to make the efficiency as high as possible, we will not work
the copper at a very high-current density, and we will keep down the saturation
of the iron. It will, therefore, be advisable to keep the pole pitch as great as 32 cms.
We will take 12 slots per pole as before, but now there will only be 4 conductors
per slot, each carrying a normal current of 333 amperes. The conductor might
be made 0-41 cm. x 1*4 cm., giving a nominal current density of about 600 amps,
per sq. cm. These will go into a slot 1 1 cm. by 4 1 cms., and we will find that with
a diameter 'of armature of 245 cms. and a length of 30 cms., after due allowance
for ventilating ducts and insulation, we can get a total cross-section of all the teeth
of 10,300 sq. cms.

The magnetic loading will be found by the formula

258 = •74x4-16x48x^^8,
^^8=1-75x108.

Dividing by 10,300 sq. cms. we get a maximum density in the teeth, B =17,000.
This is a suitable figure for a machine of this chariicter, so we may adopt the dimen-
sions given above.

It is unnecessary to go through all the steps in the calculation of the machine,
as these are very similar to those described on pages 321 and 567.

The winding of the commutating pole may consist of one turn through which
4000 amperes pass, there being two paths in parallel through the commutating
winding. In one of these halves the current will pass one way around the shaft,
and in the other the other way, to avoid magnetizing the shaft.

There will be no series winding.

Sometimes, for electrolytic purposes, it is desirable to change the voltage of the
continuous-current supply. If this must be done over a wide range and in a con-
tinuous manner, that is to say, by going from one voltage to another in infinitely

ROTARY CONVERTERS

595

sma]l steps, it is best to supply an A.O. booster for the purpose. In this cases
however, as will be seen, from Specification No. 15, it is sufficient to give a continuous
range of voltage over a small range (240 to 260 volts), and at times when a lower
or higher voltage is needed, the Purchaser is content to make the change by changing
the tapping on the transformer. This is a much more economical method than
the one requiring a booster, the copper losses in which would always be going on
whether the change in voltage were required or not. The range from 240 to 260
volts can be obtained very readily by changing the excitation and causing a reactive
drop or a reactive rise in the transformer. The theory of this method is given
below. As we only require about 4 per cent, change in voltage from the mean
in this case, it is sufficient to give the transformer a 10 per cent, reactive drop
between no load and full load (cos <^=0).

THE VARIATION OF THE VOLTAGE OF A ROTARY CONVERTER
BY THE VARIATION OF ITS EXCITATION.

A transformer having considerable magnetic leakage behaves in some respects
like a choke coil. When it is fed with a constant alternating voltage on the primary
side, the voltage at the terminals
of the secondary for a given load
will depend upon the power
factor of the load. If the current
lags, there will be an inductive
drop in the transformer; whereas
if the current leads there will be
an inductive rise. One of the
simplest ways of representing the ^
relation between the primary and
secondary voltage of a trans-
former for various loads and
power factors is that given in
Fig. 617. For the purpose of
this figure, the phase of the
secondary voltage is represented
by a vertical line, and is taken
as the standard of reference for
the phase of all other quantities.
The primary voltage will change
its phase with respect to this
datum line, but the vector OE^
representing it will always be of the same length, provided the primary voltage
remains constant. We may therefore draw the arc AE^B of & circle with its centre
at 0, to give us the locus of the point J^^. The inductive drop which occurs in
the transformer is approximately at right angles in phase to the secondary current,
and the amount of it depends upon the design of the transformer. For the purpose
of what follows here, we shall assume that a transformer can be built so as to give

Fig. 517. — Phase relations between primary O^i and secondary
OEi voltages of a transformer with current O/g lagging.

696

DYNAMO-ELECTRIC MACHINERY

any required inductive drop when carrying its full-load current at zero power factor.
Suppose, for the purpose of illustration, that a transformer is designed to give at
full-load current a reactance voltage equal to 20 per cent, of the normal voltage.
If, then, we have a full-load current lagging 30°, as shown in Fig. 517, the line
E^E^ will represent the inductive drop in the transformer, its length being 20 per
cent, of OE^, and its phase position being at right angles to the vector 01^, which
represents full-load current in the transformer. Neglecting the resistance drop,
the line OE2 is proportional to the secondary voltage, and will, with a lading
current, be less than OE^, If, now, the current leads on the secondary voltage,

as indicated in Fig. 518, E1E2, which
must still be drawn at right angles to
0/2, is so placed that OE^ is greater
than 0^1.

We see from Fig. 518 that the amount
by which OE^ exceeds OE^ depends upon
three factors : (1) The amount of induc-
tance in the transformer ; (2) the value
of the secondary current ; and (3) the
value of the angle ^g- ^^ ^^ confine
ourselves to the case where full-load
current is passing in the secondary, we
have to consider only factors (1) and (3).
We can obtain a given inductive rise of,
say, 10 per cent., either by a great induct-
ance and a small angle <f>2, or by having
a smaller inductance and a greater <^.
The amount of inductance which we
should give to a transformer in any
rotary converter installation will depend
upon certain circumstances which we
shall consider presently (see page 599).
We will assume for the moment that we
have decided upon the amount of induct-
ance : this is usually measured by the percentage reactive drop obtained in the
transformer when carrying full-load current at zero power factor.

When the load on the secondary side of the transformei consists of a rotary
converter or synchronous motor, the current can be made to lead by over-exciting
the converter or synchronous motor, and can be made to lag by under-exciting
it. In the case of a rotary converter, an adjustment of the voltage may be made
by hand by adjusting the rheostat in circuit with the shunt coils. ^Vllere the
converter is to be over-compounded, the change in the number of ampere-turns
per pole is effected automatically by the series windings. The shunt excitation
is adjusted so that at no load a lagging current is drawn from the transformer ;
then, as the load increases the excitation becomes normal, and with a further
increase in the load the converter becomes over-excited and draws a leading
current.

Fio. 518. — Phase relations between primary and
secondary voltages of a transformer when carrying
full-load current OIj, leading by the angle <^ on the
secondary voltage.

ROTARY CONVERTERS 697

The method of working out the number of ampere-turns which must be added
to the normal excitation of a rotary, in order to obtain a given rise in voltage, will
be best understood from an example worked out.

Consider a 1250-k.w., 550-volt, six-phase rotary converter, and suppose that
it is desired to make it compound from 500 volts, no load, to 550 volts, full load.
In order to proceed, it is necessary to have the following data :

1. The number of conductors in the armature and the current per conductor

at full load.

2. The percentage reactive drop in the transformer at full load.

We will assume that the machine is a 14-pole machine, like that particulars
of which are given on page 582, having 12 slots per pole and 8 conductors per slot.
We will further assume that the reactive drop in the transformer is 20 per cent,
of the normal voltage when full-load current at zero power factor is drawn from the
secondary. The amount of this reactive drop may be fixed in order to suit the
circumstances of the case, as explained on page 599.

It is best to draw the graphic diagram as if the transformer ratio were 1:1. It is
well to allow about 3 per cent, (in this case 16 volts) for ohmic drop in transformer,
armature, brushes and field windings. Thus it is necessary to generate 566 volts at
full load. Now, if the excitation of the rotary were adjusted so as to give us unity
power factor at no load, it would be necessary to over-excite the field-magnet
and make it draw a leading current sufficiently great at full load to give a rise
of 66 volts. This would cause much more heating in the armature than if we
adopted the plan of arranging the excitation so that at about half load the con-
verter is running at unity power factx)r and generating, say, 533 volts. It would
then only be necessary to over-excite it so as to obtain an additional 33 volts at
full load, and this would be done with a smaller leading component than if the
whole 66 volts had to be obtained by over-excitation. The drop in volt-age of 33
volts between half load and no load can easily be obtained by drawing a lagging
current from the transformer at no load. This is the plan usually adopted, though
it is not important that the excitation which shall give unity power factor shall
occur exactly at half load.

Let us decide, then, in the first instance, to run at about imity power factor
at 533 volts. This will fix the ratio of transformation of the transformer, which
will be so adjusted that on unity power factor we have 373 volts between rings 1
and 4 of the converter. Taking the ratio of transformation of 1 : 1 for the purpose
of our diagram, we draw the arc of the circle at a radius, OEi, of 373. At full
load it is desired to generate 565 volts c.c, that is, 394 volts a.c. between rings
1 and 4. We therefore set off 0-^2 = 394. The reactive drop in the transformer
at full load being 20 per cent, of 373, we know that the radius ^2^1 ^ equal to
74-6 ; and this can therefore be set off, taking E2 as the centre. We thus obtain
the position of J^^, which must lie on the arc of the circle. We now know that the
secondary current OI2 is at right angles to E^E^ ; its phase position is therefore
ascertained. The full-load working current 01 to in phase with the secondary
voltage being known (in this case 1180 amperes per phase), we can at once set
off Iwlzj the leading wattless current at full load (in this case 388 amperes), and

698

DYNAMO-ELECTRIC MACHINERY

obtain the length of O/g (1220 amperes), the full-load current for which the trans-
former must be designed. We can now proceed to find by how much the field-
winding of the rotary must be over-excited in order to cause the leading wattless
current I^Iy, to flow. This depends upon the number of conductors in the
armature.

If each pole of a rotary converter of normal pole pitch is over-excited by an
amount equal to 0-87 of the normal c.c. armature ampere-turns, a leading wattless
current of full-load value will flow in the armature. In this case the continuous
current per conductor will be 168 amperes, and there being 48 turns per pole,
the continuous-current ampere-turns per pole will be 168x48=8064.

Fig.' 519. — ^Phaae relations at half load.

0 la

Fig. 520. — Phase relations at no load.

From Fig. 518 we see that l^l^ is equal to 0-33 of 0/^, so that the amount of
over-excitation required to make Itol^ flow will be equal to 0-33 x 0-87 x 8064 = 2340
ampere-turns.

Before we can settle the right number of series turns to put on the poles of the
rotary, we must consider the conditions of running at half load and no load. Fig.
519 shows the clock diagram at half load. O/u^ is now half the value it had in Fig.
518. The excitation of the pole is now nearly normal, so that 01^ is almost in phase
with OE^. The reactive voltage in the transformer E^E^ is now only 36 volts, on
account of its phase position 0Ei=0E2, so that the rotary has a generated voltage
of 533 and a terminal voltage of 525.

The clock diagram at no load is given on Fig. 520. Here the reactive voltage
E^Ei is in phase with 0^2> ^^ w® ^^^ calculate how much it should be in order to
get the desired voltage drop at no load. To get 33 volts c.c. we require 23-3 volts
A.c. Now, if 1220 amperes give a reactive voltage of 74-5, it will take 383 amperes
lagging current to give a drop of 23-3 volts. At no load, therefore, 01^ must be a
lagging current of 383 amperes per phase. The amount by which the excitation
must be below normal to cause this lagging current to flow is calculated as follows.

ROTARY CONVERTERS 599

Full-load current wattless ( ^ 1180 amperes) requires 8064 x 0-87 = 7000 ampere-turns

per pole ; therefore

383
7000 X :pr^ = 2280 ampere-turns
lloU

below normal excitation are required to niake 383 amperes lagging current
flow.

Suppose that the normal excitation for 500 volts is 3643. Then the shunt
excitation would be adjusted to 3643-2280 = 1363 ampere-turns per pole at no
load. At 550 volts this would increase to 1500. If, now, the proper excitation for
550 volts unity power factor is 4500, we must have 3000 + 2340 = 5340 series ampere-
turns in order to draw 388 leading amperes through the transformer. As the full-
load current is 2360, we require 5340 -h 2360 or about 2-5 turns per pole in the series
winding. In the above we have assumed that the reactance voltage follows a straight
line law ; in other words, that it is proportional to the load on the secondary. This
is not always true, particularly in transformers containing " reactive " iron. When
we cannot assume true proportionality, a curve showing the reactive voltage at
different loads should be obtained from the designer of the transformer, and this
can then be worked to in setting out the reactive voltage E^E^ in the graphic
construction. In actual practice, in the case considered, one would put three series
turns per pole on the 1250 rotary considered above in order to be on the safe side.
It is then an easy matter to divert some of the series winding to obtain the amount
of compounding that we require.

At half load we would have 5340-5-2 = 2670 series ampere-turns and 1430 shunt
ampere-turns. As the total 4100 is just a little more than the normal excitation
at 525 volts, the current would lead just a very little and give a clock diagram as
shown in Fig. 519.

Power factor on the H.T. side and L.T. side. It will be seen from Figs. 517 and
518 that where the transformer has considerable reactance the angle ^2 between
the current and the voltage OE^ may be either greater or less than the angle <t>i be-
tween the current and OEi, When the current lags <l>^ is less, and when it leads
<f>2 is greater than </>i. Now, the amount of heating of the copper on the armature
(see p. 542) depends upon the value of ^2- Where the transformer reactance is very
great, <l>i may be nearly zero at full load, though ^2 ^*s a high value. Thus the
power factor on the h.t. side is nearly unity, and yet there may be such a low
power factor on the l.t. side that considerable heating occurs in the armature
copper.

In cases where it is desired to run the rotary plant on a leading power factor
(that is to say, leading on the h.t. side), it is desirable to keep the reactance voltage
of the transformer fairly low ; and if a wide range of compounding is required, it
is best to carry it out by means of a booster, so that the rotary has only to supply
the leading current required on the h.t. side, and not a heavy additional leading
current to compensate for the reactive drop in the transformers.

600

DYNAMO-ELECTRIC MACHINERY

SPECIAL PRECAUTIONS NECESSARY WHEN THE FREQUENCY IS

UNSTEADY.

On pages 337 to 356 we considered the precautions which
should be taken to ensure the good parallel running of syn-
chronous machines. The same rules apply to the case of
rotary converters, but here the disturbance is not so often
in the converter itself as in the prime mover supplying the
alternating current. In all cases where there is reason to
believe that the frequency of the system is not constant,
particulars should be obtained of the nature of the disturb-
ance, and care should be taken to see that the natural period
of phase-swinging of the converter does not correspond with
the period of the disturbance ; otherwise resonance may be set
up, which, if it does not prevent parallel running altogether,
may seriously aifect the conmiutation. In one case within
the experience of the author, a rotary converter would not
run in parallel with slow-speed engine sets because its natural
period of phase-swing coincided almost exactly with the period
of the disturbance. A tachograph record of the speed of the
converter was taken during the interval of time between
the instant of switching the converter to the bus-bars and
the instant at which it broke step. Fig. 521 gives the record.
It will be seen that the speed was fairly constant during the
first second or two. Then the phase-swinging was augmented
and then diminished as the natural swing of the machine got
into and out of step with the disturbance. A change in the
excitation was sufficient to give almost perfect resonance, and
the convert-er then broke step. The addition of considerable
self-induction in circuit with the slip-rings was foimd to so
alter the value of /„ (page 339) that parallel running became
possible without any change being made on the dampers.
On the other hand, if very heavy dampers had been added to
the pole shoes the machine would have run even imder con-
ditions of perfect resonance.

For the satisfactory operation of a rotary converter running
on a circuit with an unsteady frequency, it is not sufficient
merely to reduce the phase-swinging to a point that makes
parallel running possible. It is necessary to reduce it so that
it no longer interferes with the commutation.

From what was said on page 338, it will be understood that
when a converter is phase-swinging, it carries a motor load
when the armature is being accelerated, and it carries a
generator load when the armature is being decelerated. These
motor and generator loads, being uncompensated, produced a
field distortion that may very seriously interfere with the

ROTARY CONVERTERS 601

excitation of the commutating pole. The designer should be able in any given
case to work out roughly the amount of distortion produced.

A case is worked out below numerically. The actual amoimt of the phase-swing
depends upon the efiectiveness of the damper. On page 352 we saw that the
efiectiveness of a damper can be conveniently expressed in terms of the slip which
the machine would have if run as an induction motor with the damper acting as a
squirrel cage. A case was worked out showing how the effectiveness of a damper can
be roughly estimated. We shall employ the same notation as on pages 339 and 354.
By the symbol s we denote the slip there would be on the machine if run at full
load as an induction motor using the damper as a squirrel-cage winding. Thus,
if the slip would be 2 per cent., 8 =0 02. Then 2irns is the angular velocity of the

slip in a two-pole machine, and is the angular velocity of the slip on a machine

having p pairs of poles. Now, if an angular velocity of gives the full-load torque,

-- (see page 354),

9-81 X i2^ X 2-

where E is the voltage and / the current measured on the continuous-current side
of the converter, then, for any relative angular velocity (d - x), the torque exerted
by the damper will be

h(a-x) =jrH= — rfi — s ^ — kilograms at a metre radius.

^ ' 981 X jBp, X 2;r X 2r»5 ®

Now the torque required to produce an angular acceleration a is

aa =7r^:ra Idlograms at a metre radius,

and the synchronizing torque for a relative angular displacement (a - x) is

, . EIx/3xp . . , «.-.

''<°-^)=981xii.xV<"-^^ (see page 341).

Here a is the angular displacement of the armature and x is the angular displace-
ment of the vector representing the supply voltage relatively to a uniformly rotating
vector. We may take x=^A sin cu^, where A is the amplitude of displacement of
the voltage vector.

In what follows, therefore, we may use the contractions.

a =

9-81'

EIxp

9-81x2^,x2irx2;rrw'
, EIx/Sxp

a = —

602 DYNAMO-ELECTRIC MACHINERY

THE PHASE-SWINGING OF SYNCHRONOUS MOTORS AND ROTARY
CONVERTERS. WITH AND WITHOUT DAMPERS.

The problem of damping * of the phase-swing of a synchronous motor or rotary

converter is rather different from that of the damping of oscillations of a generator

set np by irregularities in the turning moment, because the torque due to the damper

is not proportional to the angular velocity of the phasenswing u, but to (d-x),

where i is the rate of change of the angular displacement of the phase of the impressed

voltage from a voltage of constant frequency.

We get (see page 601)

aa + 6(d-i) + c(a-x) =0 (1)

li x=A sin (o^, the solution of the above equation is

-^^===.sm(o.-fe + tan ^j) (2)

where 6 = tan"* — , oi =2'trnd, fid is the frequency of the phase-swing, and k =(a(i>* - c).

If b =0, that is, if there is no damping, a = - —^ sin orf, and as k =aw2(l - q)^

a = - - — ^ sin tot, where q = — ^i
l-q ^ au>*

That is to say, the original swing is multiplied by —2—. Wbere y=l, a will
become infinite, if 6 =0 (see page 340). ^

It is interesting to enquire how great the swing will be where ^=1/ and the
damping is such as one would find in practice. Writing k =0 in (2), we get

A>/c^ + (i>6* . / ^

a = sm { cut + 6 +

Now let us give values to a, b and c such as we might have in the 1250 K.w.
rotary described on page 570, taking hist of all a 3 per cent, damper (see page 354).

a=1526. ^-_ l-25xlO«x7 -2ixlQ--

9-81' 9-81 X 715 X 6-28 X 6-28 X 50 X 03 "■'

1-25 xlO«x 1-4x7 „^Q inn 1
0 = 9-81 X 7-15x6 28" " ^ kilograms at a metre.

Here we have put ^ = 1*4. This is worked out in the method given on
page 294.

The value of w which will make A; =0 is w =16'4. That is, rid =2*6.
Then (d6 =164 x 21 x 10» =344 x 10*. J^^H^b^ =4*4 x 10*.

a = -1-28 ^sinfwi + tan"^— +|j,

tt = 268 X 1-28^ sin ( arf + tan"!

urf+tan"^ — + ^j.

♦ " The Function of Damping-coils in the Parallel Running of Alternators," I. DOry, Elekiral
«. Maschinenbau, 27, p. 315, 1909 ; " Theory of Damping in Parallel Running," C. F. Guilbert,
Lumiere Electr., 9, pp. 355 and 387, 1910 ; " The Proportioning of Amortisseurs," Fimde, Elecirot.
u. Maschinenbau, 27, p. 1073, 1910; "The Amortisseur Winding," M. C. Smith, Oen. Eltct. Bev^.
16, p. 232, 1913.

ROTARY CONVERTERS

603

The maximum value of aa =3*6 x 10* x ^ kilograms at a metre.

The maximum value of ha =2*1 x 10* x 16*4 x 1-28 x A =44 x 10^-4.

The maximum value of ca =36 x 10* x ^.

The maximum value of the synchronizing torque is obtained by finding the
maximum value of c(a^a:). The quantities take up the phase positions shown in
Fig. 522, from which it will be seen that

c(a-x) ='8cx, a-a;=*8a?.

Now, how big may we expect to find x in ordinary practice ? If a 50-cycle
generator, supplying the network on which the rotary runs, having 64 poles, running

- Fio. 522. — Phase positions of the various torques exerted on a damped synchronous motor

when running on a circuit of varying frequency.

at 1*57 revs, per sec, has an angular irregularity of ^^yth, we will have an angular
variation of '004 x 157 x 2ir radians per second. This will give rise to an angular

displacement of the phase of „ ^-^ = 00242 radian on the generator, or

•00242 X f J = Oil radian on the converter. So the maximum value of a; = Oil =^.
But a-x= '9>z. Therefore the maximum synchronizing torque under the conditions

will be

c(a-x)=,

J^/x 1-4x7

x-Sx-Oll.

9-81 X jRp, X 27r

That is to say, the synchronizing torque may be about 1*4x7 x '8 x "Oil
of full-load torque, or 8^ per cent. This would be quite enough to affect the
commutation.

If we make the damper of sufficient conductivity to reduce the slip to 1^ per
cent, at full load, we reduce the value of c(a -a;) to 3-8 per cent, of full-load torque.

That is to say, that the number of ampere-turns applied by the armature to
the commutating pole in exc-ess of the correct number at one part of the swing and
subtracted from the commutating-pole excitation at another part of the swing is
not more than 038 x 7000 =265 (see page 599), and as the total number of effective
ampere-turns on the pole is 4720, the interference with the commutation is only
very slight.

604 DYNAMO-ELECTRIC MACHINERY

We see, therefore, that if we fit very good dampers and have the commutatizig
conditions so that the commutation is not sensitive, we can run satisfactorily
even if A; =0.

SMALL ROTARY CONVERTERS.

The specification for a small converter should be as simple as possible, so as to
enable a manufacturer to quote on his standard plant, and should be confined
to a statement of the characteristics required, such as given in Clause 242, page 584,
a statement of the purpose for which it is required, and only such special clauses
as are absolutely necessary. The starting of small converters (up to 100 K.W.
in small systems and up to 300 k.w. in large systems) is generally carried out by
taps on the transformer without any starting motor, as shown in Fig. 507. For
500-volt machines, 3-phases are more common than 6-phase for small machines,
because the extra complication in wiring and slip-rings is hardly worth while
when the current per ring is very low. When running from c.c. to a.o. it is also
usual to do without an exciter (see page 559), when the cost of this would be an
excessive fraction of the total cost of the plant. A series winding will in most
cases be sufficient to keep the speed of the converter within reasonable limits.

CHAPTER XX.

PHASE ADVANCERS.

The phase advancer stands in the same relation to an induction motor as an
exciter does to a synchronous motor. The exciter supplies a continuous current to
magnetize the field-magnet, so that the power factor of the motor may be kept
near unity. The phase advancer supplies a current slowly alternating with the
frequency of the slip, which acts as a magnetizing current and obviates the neces-
sity of any magnetizing current being supplied to the stator ; and thus the power
factor of the motor is improved. Where it is desired to run an induction motor on
a leading power factor, the phase advancer is made to supply a magnetizing current
greater than the normal magnetizing current, and it is then found that the stator
draws a leading current from the line, just as a synchronous motor does when it is
over-excited.

A phase advancer can, of course, only be used in conjunction with a motor
which has a wound rotor with the ends of the windings brought out to slip-rings.
Squirrel-cage motors cannot, therefore, be used with phase advancers.

It is only when the user of the motor has some interest in the power factor
that he will go to the expense of installing a phase advancer. When power is charged
for at so much per kilowatt-hour, independently of the power factor, the user of
the motor will prefer to draw his magnetizing current from the line free of cost.
But where an extra rate is charged for wattless current, or where a rebate is made
for power taken at a good power factor (and it seems probable that in the future
such systems of charging will be more common), it may be worth while for a user
to supply his own magnetizing current. There are cases, too, where the user
generates his own power, as, for instance, where a Corporation has large induction
motor-generator sets on its mains. Here phase advancers could often be installed
with advantage.

It will be chiefly in connection with large motors that these machines will
be installed, because the cost of the extra appliances is too great a proportion of the
cost of the motor where the motor is small.

In deciding whether or not it is worth while to install a phase advancer, the
following matters should be taken into account :

(1) What monetary advantage is to be gained by improving the power factor
of the motor ?

606 DYNAMO-ELECTRIC MACHINERY

(2) What will be the cost of the advancer and switch gear, and what extra

power will it consume ?

(3) What extra attendance will be required ?

(4) What improved performance can be obtained from the motor ?

The strongest cases for the use of a phase advancer are those where the mains
or the plant in the power house are loaded to their utmost, and it is desired to
install more motors. In such cases the addition of a phase advancer of 10 k.v.a.
capacity to one or two big motors may liberate some hundreds of k.v.a. to be used
elsewhere, and obviate the necessity of extensive additions to the plant. The
advantage which the advancer has in these cases arises from the fact that it supplies
the magnetizing current at such a low frequency and low voltage that the total k.v.a.
supplied is only a very small fraction of the wattless k.v.a. which would otherwise
have to be supplied at the frequency and voltage of the station. The k.v.a. supplied
by the phase advancer bears the same proportion to the k.v.a. saved in the stator
as the slip of the motor bears to the synchronous speed.

The monetary advantage to be gained by improving the power factor can
easily be found where the power is supplied by a public company, and definite rates
are charged for true power and for wattless current. Where the user supplies his
own power, it is not so easy to arrive at the cost of the wattless current. It would
be necessary to make a close enquiry into the circumstances of each particular case.
From a number of investigations made by Prof. Arno and Mr. Conti, engineers
to certain Italian power companies, with respect to working cost of generation and
transmission, as affected by the power factor of customers, it was found that the
average cost was almost proportional to

^EIco8<f> + ^EI*

so that perfectly wattless k.v.a. would cost one-third of the same k.v.a. at unity
power factor.

When a power company has a large induction motor-generator, such as that
described on page 448, running on its system, it would be possible, by means of a
phase advancer, to run it on a leading power factor, so as to compensate for the bad
power factor of other motors running on the same system. Instead of running at
0-88 power factor and drawing 640 k.v.a. wattless from the line, it might be made
to run at a power factor of, say, 0-95 leading, and supply 460 leading wattless k.v.a.
to the line, making a total change of 1000 wattless k.v.a. This would be sufficient to
change a total load of 1660 k.v.a. at 0-8 power factor into a load of 1330 k.w.
at unity power factor. If the cost of generation and transmission were as found
by Arno, viz. proportional to ^EIcoa<f> + iEI, the cost of a year's run of 3000
hours at O-bd. per unit would be

{§ X 1330 + ^ X 1660) X 3000 X 0-5 =£9000

without the phase advancer, and

1330x3000x0-5 ^^o^n
240 =^^^^

with the advancer, giving a difference of £650. The total cost of a suitable advancer
of 30 ^. V. A. capacity, including the cost of extra copper in the rotor of the induction

* See Prof. Gisbert Kapp, Inst. EUc, Engineers' Joum., vol. 50, page 351.

PHASE ADVANCERS 607

motor, would not be more than £250. The losses in the phase advancer would
be about 6-5 K.w. (see page 570), and the saving in the efficiency of the induction
motor 2-5 K.w., giving a total loss of 4 k.w., costing £25 per annum. If we esti-
mate the extra attendance at £10 per annum, we make a saving of £615 per annum
for a capital outlay of £250. This is on the basis of Amo's figures for the cost of
wattless K.V.A., and the calculation would have to be modified to meet the cir-
cumstances of any particular case. Where the motor is not so large, the saving
is not so great ; but even down to sizes of 200 h.p. circumstances may be such as to
make the addition of an advanc-er well worth while. At present the prices of these
machines are very high, because the manufacturer regards them as special machines ;
but in the future the price will no doubt be very much reduced, and then smaller
motors can be fitted economically.

In cases where a new motor is being supplied together with a phase advancer,
the specification of the motor may vary in some respects from the specification of
a motor taking its magnetizing current from the line. Specification No. 7a gives
particulars of the variation which might be made in Specification No. 7 (page 438)
to suit the case where a phase advancer is to be supplied with the motor.

608 DYNAMO-ELECTRIC MACHINERY

SPECIFICATION NO. 7a.

1600 H.P. 3-PHASE INDUCTION MOTOR, 3000 VOLTS, 60 CYCLES,
246 RP.M., INTENDED TO BE RUN ON LEADING POWER FACTOR.

This specification will contain all the Clauses in Specification No. 7,
p. 438, with the following exceptions and additions :

Instead of Clauses 88 and 89, substitute the following :

characterutics 300. The motor shall have the foUowing characteristics :

of Motor ^

with PhasA

Advanc6r. NoFmal output 1500 H.P.

Normal voltage at ter-
minals 3000 volts.
Frequency 50 cycles.
Number of phases 3.
Speed 246 revs, per minute.
Power factor at full

load 0*95 leading.

K.v.A. at no load 400 leading.

How connected to load Direct connected through flange

coupling.
How connected to

phase advancer Belted.

Size of pulley on motor

shaft 37 inches in diameter.

10 inches on face.
Temperature rise after
6 hours full load

run 40° C, by thermometer.

Over load 25 per cent, for 3 hours.

Temperature rise after
3 hours 25 per cent.
. over load 55° C. by thermometer.

Maximum torque 5 times full-load torque.

Puncture test 6600 volts alternating at 50

cycles appUed for 1 minute
between stator windings and
frame.
4000 volts alternating at 50
cycles for 1 minute between
rotor windings and frame.

PHASE ADVANCERS 609

301. The contract includes the delivery of the motor and Bxtentorwork.
phase advancer at the sub-station of the Corporation, together

with bedplates, bearings and pedestals, and the erecting,
aligning and coupling of the same to the 1000 K.w. generator,
and the lining up and belting of the phase advancer. The
switch gear and starting gear are provided for under another
specification.

302. After the rotor bars have been connected together. Testa on Eotor
but before they are connected in star or in mesh, a pressure of ^^^'
5000 volts shall be applied for 1 minute between phases A

and B, B and C, and C and A. The phases shall then be
connected in star or in mesh, and a pressure of 4000 volts shall
be appUed between copper and iron. This latter test shall be
repeated after the motor has been run at ftdl load for 6 hours,
and while it is still hot.

Id addition to Clause 100, there should be the following paragraph :

303. The tender shall state what further losses are Efficiency.
occasioned in the motor when it is running in conjunction

with the phase advancer under the conditions specified in
Clause , and also the losses in the phase advancer itself and
its connections. In the above statement the contractor is
entitled to take credit for any diminution of losses in the
stator winding or elsewhere.

Clause No. 102 will be modified by changing paragraph (3), which
should read :

304. (3) Power factor test. During the temperature run Tests on site.
the phase advancer shall be in circuit with the rotor winding,

and the power factor of the combination shall be measured
by means of a power-factor meter, and also by the two-
wattmeter method. The motor shall also be run with the
CO. generator unloaded, and the power factor again
observed, to see that the conditions described in Clause
300 have been met.

The Clause as to brush gear will be the same as Clause 106, except
that there should be added the following :

305. The slip-rings and brush gear shall be designed to Brush Gear.
carry the full-load current continuously without excessive
heating or excessive wear.

W.M. 2 Q

610 DYNAMO-ELECTRIC MACHINERY

In specifying a phase advancer, the following particulars" should be given:

(a) The type of motor to which it is to be fitted.

(b) The voltage and frequency of supply.

direction, or whether it reverses or stops often.

(d) The stand-still voltage of the rotor and the number of phases,
(c) The full-load working current.

(/) The power factor of the motor at full load and the no-load'current.

ig) To what value it is intended to alter the power factor ; or, in other words,

what total change in wattless k.v.a. is desired.
(A) Particulars of the slip-rings and brush gear on the rotor, and their current -

carrying capacity on continuous service,
(i) The method of starting the motor.
(J) The proposed method of driving the phase advancer. With high-speed

motors it may be direct connected ; with slow-speed motors it may be

driven by a belt or connected to some other running machinery, or to an

independent motor.

The following model specification shows how such particulars might be given :

SPECIFICATION NO. 16.

30 K.V.A. PHASE ADVANCER.

Rating of 306. The Phase Advancer is to be suitable for putting in

Motor. •••11 A 1 11

cu-cuit with the rotor of a 1500-h.p., three-phase, 3000-volt
induction motor, running at 246 r.p.m.

SStor?' ^^'^' '^^^ motor will be direct-connected to a 1000-k.w.

c.c. generator which will supply a power and lighting load.
The motor will run 18 hours a day for the greater part of
the year, with a load which will vary according to the con-
ditions of service. The over load on the motor may amount
to 25 per cent, for 2 hours.

PowerFactor 308. Thc Calculated power factor of the motor is 0*88 at
full load, and the no-load current about 90 amperes.

Desired Power 308a. It is dcsircd that the motor shall be capable of
yielding 400 leading wattless k.v.a. at all loads.

standatiu 309. The rotor is woimd for three phases, and the standstill

voltage will be 1460 volts per phase.

siipat Full 310, The slip at full load will be about Ij per cent.

PHASE ADVANCERS 611

311. The full-load working current of the rotor will bewOTking

rfc P^f\ Current.

250 amperes.

312. The slip-rings and brush gear of the rotor are designed sup Rings
to carry 400 amperes per ring in continuous service. ^r^"**

313. The motor will be started upon a water resistance, Method of
while the phase advancer is cut out. After the resistance °**
has been short circuited, the phase advancer will be thrown

into the circuit by means of a double-throw switch.

314. It is proposed to belt the phase advancer to the Driving Power,
motor.*

315. The above particulars are given for the guidance of General
the Contractor, but it is not intended that the amount of the °'°^* °°*
leading wattless k.v.a. of the motor shall be limited to exactly

400 K.V.A. , and so long as no overheating occurs, the wattless
K.V.A. may with advantage be increased.

316. The brush-gear and commutator shall be amply Bmsh Gear and

J • J J. x' 1 •j.i_ J. "L J-* Commutator.

designed so as to run contmuously without overheating.

317. The temperature rise after 6 hours' full-load run shall Temperature
not be more than 45° C. by thermometer.

318. The phase advancer shall be subjected to a testing Puncture Teat,
pressure of 100 volts (alternating) applied for one minute
between armature and frame and field coils and frame.

319. With the phase advancer the Contractor shall supply Bedplate,
a bedplate adapted for bolting to the bedplate of the driving
motor, which is already existing, and particulars of which will

be supplied.

320. Foundations will be supplied by the Purchaser to Foundations.
templates furnished by the Contractor. Cables from the cawes.
motor to the throw-over switch, and from the throw-over
switch to the advancer, will be supplied by the Purchaser,

the throw-over switch being supplied umder another contract.

321. The contract will include the making of all proper setting to
connections and the setting to work of the complete plant. ^°'*^*

* In the case of high-speed motors it is convenient to direct-connect the phase
advancer.

612

Losses.

Oil-throwing.

DYNAMO-ELECTRIC MACHINERY

322. The Contractor shall state the amount of the losses in
the phase advancer when operating under the above specified
conditions at full load.

323. The journals, bearings and housings of the advancer
shall be designed so as to be perfectly free from oil-throwing.

THE DESIGN OP A 30-K.V.A PHASE ADVANCER.

For the general theory of phase-advancing the reader is referred to the specifica-
tions and articles* quoted below.

We propose here to give the method of designing a phase advancer to meet
certain conditions of service. We will take the conditions set out in Specification
No. 16.

The armature may either be of the open-circuit star type (see Journal of the
Institution of Electrical Engineers , vol. 42, p. 612, Fig. 10, 1909), or of the closed-
circuit type. Both kinds of armature commutate well. The first (see Fig. 523)

Fia. 523. — Diagram of open-circuit armature
with several branches in parallel under wide
bruflh belonging to each phase (see Journal of the
IrutUuHon of Electrical EndneerSt vol. 42, p. 612,
Fig. 10, 1909.)

Fig. 524. — Closed-circuit armature forming a
mesh connection between the phases.

is suitable when the current to be collected on the commutator is very great and
the voltage to be generated is small, say not more than 15 volts. It enables a very
wide brush (extending over 0-7 of the pole pitch) to be used. The second type

♦ Brit. Pat. Specification, No. 15,470 of 1896.

M. Walker, " The Improvement of Power Factor on Alternating-current Systemn,*' Joum.
Inst, Elec. Engineers, vol. 42, page 599 ; also ibid. vol. 50, page 329.

'' Improving the Power Factor of Induction Motors," Elec. Engineering, 6, p. 229, 1910 ;
Electrician, 64, p. 1064, 1910 ; " Phase Compensation of Induction Motors," Brit. Pat. 28,383
(1911), Engineer, 114, p. 507, 1912 ; " New Machine for Phase -compensation of Single- or Poly-
phase Induction Motors," A. Scherbius, Elektrotech. Zeitschr., 33, p. 1079, 1912.

Dr. G. Kapp, " On Phase-Advancers for Non-synchronous Macl^nes," Electrician, vol. 69,
pages 222, 272.

" Improvement of Power Factor in Power Systems," Bauer, Schtpeiz. elektrot. F«rfiii,^Bull. 4.

p. 304, 1913 ; " Theory and Applications of the Leblanc Exciter," Ehrmann, Lumiere Eled., 22,
p. 291, 1913 ; " Improvement of Power Factor." Kapp. Elektrot. Zeit,, 34, p. 931, 1913 ; " Phase
Compensation," Fynn, Elec, World, 62, pp. 28, 75 and 132, 1913 ; " Phase Variator." Campos,
AUi deW Assoc, EleUr, Ital., 17, p. 221, 1913.

PHASE ADVANCERS

613

(see Fig. 524) is suitable when the current is not very great and the voltage is higher.
As the current, in the present case, will only be a little over 300 amperes and the
voltage to be generated will be about 70, we will choose the mesh-connected type.

In this case the rotor has a three-phase star-connected winding having a stand-
still pressure of 1450 volts per phase. The working current (that is to say, the
current in phase with the voltage) will be about 260 amperes, which can be collected
on a comparatively small collector. To find the rotor current necessary to make
the motor run at 0-95 leading power factor, proceed as follows :

Set ofE a vertical line representing 260 amperes, as shown in Fig. 525. The
power factor of the motor is 0-88, so that without the advancer one would have a
lagging current equal to 47 per cent, of the working current. If the advancer
caused the rotor to take a leading current of 47 per cent, (that is, 122 amperes),

Leading Current

Fig. 525. — Construction for finding valae of rotor current required to produce a given leading

power factor.

the power factor at the stator terminals would be nearly unity. If, now, it is desired
to make the power factor at the stator terminals 0-95 leading we must supply to
the rotor an additional 31 per cent, of leading current, making 202 amperes watt-
less in all. Adding as vectors the 202 amperes wattless to the 260 amperes working
current, we get 330 amperes per phase for the rotor when nmning under these
conditions. This is the current for which the advancer must be designed.

Next, as to the voltage to be generated by the advancer. As the armature of
the advancer is to be mesh-connected, it is simpler to take the voltages across the
slip-rings than the voltage per phase of the star winding. Indeed, as the motor
would work the same whether it were mesh-connected or star-connected, we may,
if we like, consider it mesh-connected, as we have done in Fig. 527. If the normal
slip of the motor at full load be 1-25 per cent., the e.m.f. generated by the slip
will be 31 volts measured between rings. Lay ofE as in Fig, 526 the vertical line
OEa to represent this voltage generated by the slip in phase A. In Fig. 525 we have
found the angle by which the current must lead on this voltage, so we can set off

614

DYNAMO-ELECTRIC MACHINERY

the line Oa to represent the current in phase A (see Fig. 526). Similarly Ob and Oc
represent the currents in the other phases. We should allow about 7 volts for

Fig. 520. — Construction for finding the voltage required to be generated by advancer.

pressure drop in brushes and in the resistance of the advancer. This will be repre-
sented by EaR in phase with Oa. Then there will be some reactive drop in the
field coils of the advancer. We may provisionally allow 6 volts for this, and after

the machine is calculated we can make a check
calculation to see if it is enough. This is repre-
sented by RX, There is no reactive drop in the
armature, because the compensating winding
wipes out its field. We see that, if we add a
voltage XV, parallel to ha, we shall get a re-
sultant voltage OF in phase with Oa ; and this
is what we want. If, therefore, we excite the
advancer with a current which is in phase with
the sum of Oa and - Ob (shown by the dotted
line 6a), we can make the current lead by the
right amount. The voltage to be generated by
the advancer is therefore given by XV, which,
when scaled off, gives us 49*6 volts. It will be
seen that the projection of OV on the vertical
line gives us OVr, which is greater than OE^,
If this voltage OFr is greater than is necessary to drive the working current
through the rotor circuit, the only effect will be that the slip of the rotor will be
reduced until we get the right working current for the load. If it should prove

Fio. 527. — Diagram of mesh-connected
phase-advancer armature a, 6, e, field con-
nections P. Q and R^ and mosh-oonnected
rotor A,B,C of induction motor.

PHASE ADVANCERS

615

that OVr is not sufficient to drive the working current, then the slip of the
motor will be increased.

From Fig. 526 it appears that with 49-6 volts generated by the advancer the
slip will be slightly reduced. We thus arrive at the rating of the advancer, namely,
49*6 volts between terminals and 330 amperes per phase.

We have next to decide how the advancer shall be excited. If the machine be
excited by means of a series winding, it will have the general characteristics of a
series-wound c.o. generator ; that is to say, if the speed is high enough to make
it excite when connected in circuit with a given resistance, it will immediately
take a load sufficient to saturate the iron of the magnetic circuit, and the load will
only be limited by the state of saturation. So that whether the induction motor
is on load or not, the voltage generated by the phase advancer will be fairly constant,

i-o

-o*9

o-d

oo-2r

0-6
30.5

So-s

^ -

OO-I

0^^^^^^ a^H^Kaa^ a^^^m^^^m ^^^^^^^ ^^^^^^^ m^^^^^mi m^^^m^^^ i^^i^^HB.^ — ^^^^^^ _^i^^iMM«. i^^^h^b^mi ^^^^^^mm

^^^^^^ ^^^^^0m i^^^^^"" ^^^^1^^

^^^^^^^tm i^^^H^BM ■^^■^^^iSi^— ^-^^^" ^^^^^t^mmi ^m^h^b^^ ^maim^^^^m ^^^m^i^^^ ^^^■■^^hm i^^^^h^b^m i^^^^^^^^ ^^mmi^^bm

o-i o-z 0-3 (Hb 0-5 0-6 o^ o^ 0-9 10

LOAD a6 A fraction of full toad

FIG. 528.

I'l Z-2

depending on the speed, and the flux which saturates the frame. When the motor
is loaded, the e.m.f. generated in the rotor bars helps to increase the rotor current,
and thus brings about a little more saturation of the frame ; but this is only sufficient
to make an unimportant increase in the leading current drawn by the stator from
the line. The relation between the leading wattless k.v.a. and the load on the
motor will, in fact, be as indicated in Pig. 528. The point at which the curve cuts
the no-load line will depend upon the speed of the advancer and the resistance
in series with it. In order to make the curve strike the desired point, 0-37 in our
case, it is necessary to see that with 330 amperes flowing round the field coils we
are generating the desired voltage.

As the characteristic given in Fig. 528 is the one which we desire, we will adopt
a series winding in this case. If it had been necessary to control the power
factor within narrow limits at all loads, we should find a separately-excited phase
advancer would be more suitable.

616 DYNAMO-ELECTRIC MACHINERY

Theoretically, three salient poles (equivalent to two magnetic poles) are quite
enough for a machine of the rating required in this case, but a machine of six poles
(equivalent to four poles magnetically) is more likely to fit in with standard frames
and standard punchings. We will therefore decide on six poles. This will give us
six brush arms, two in parallel in each phase. There will be 165 amperes per brush
arm, and 165 4- 1-73 = 96 amperes per conductor.

It will not be worth while to cut down a machine of this type to the smallest
possible size, because the addition of a little superfluous material will not increase
the cost by a very large percentage, and when we are making a machine we might
as well make it so that without much further development it may be used in a
large variety of cases. If we take a large L^L constant of 10 x 10^ cubic cms., it
will not be excessive, though very ample. A diameter of 46 cms. is suitable for
a speed of 750 revolutions per minute, and the length of iron may be 19 cms.

The easiest way of designing a phase advancer of this type is to proceed as if
it were a continuous-current machine whose voltage is 1 41 times greater than the
virtual voltage called for in the specification. The armature need not differ in any
particular from a continuous-current armature. The field winding will be provided
with series exciting coils and compensating windings connected to the various
phases in the manner described below.

The main points to look to, that are not found in a continuous-current design,
are :

1. The machine, though having six salient poles, is a 4-pole machine magnetically,

and we must remember this when fixing the dimensions of the iron behind
the slots.

2. The voltage to be generated as a continuous-current machine is 141 times

greater than the virtual voltage called for.

3. The fluxes in the salient poles which constitute magnetically a pole-pair

are 120^ apart in phase, so that the voltage generated in an armature
coil, which lies partly under one pole and partly under another, is only
0-86 of the voltage that would be generated if the two poles were carrying
the maximum flux at the same time.

4. It is necessary to arrange the series winding on each pole so as to cause the

flux to be the right amount ahead in phase of the current carried by the
armature conductors passing under the pole.

5. It is desirable to arrange the compensating winding so that its effect is equal

and opposite to the armature winding adjacent to it, and for this purpose
it is necessary to have regard to the phases of the currents in the armature
and field.

6. It is desirable to provide a commutating flux which shall be proportional

to, and in phase with, the current to be commutated.

The current loading. We begin, then, just as we would on a continuous-current
generator. The voltage to be generated is 49-6x141=70 volts. There are six
ways through the armature, each carrying 96 amperes. If we choose 72 slots with
4 conductors per slot, we get 288 conductors, and these multiplied by 96 give us
27,500 ampere-wires, a fairly easy current-rating for an armature 46 cms. in diameter.

PHASE ADVANCERS

617

Phase Advancer. ,^

Date ^^^-9^.19/3.. Type CCN .«¥•! MOTOMh ItOTAIIY <? .Pole. =4. .El«c Spec.../P..

KV.A.2.9. ; P.F.r^.; PhMe<? : Volts «;?^;yK?.^?faX.: Amps per ter.3^<?....: Cycles .:^^.; R.PM.7^P...: Rotor Amps..

H.P. Amps pu cood. ^^.. . Amps p. br. ann. /<^<5. Temp rise .^0 C. Regulation Overload

Customer : Order No.

Quot. No. ; Perf. Spec-

Fly-wheel effect

F««e#^5, Cira«a/45 ;GapAreai»7'.6<?;r"_
Air - Ag ■

Ac B ; poss. laZa

■ ; l«Z«

Circum.

190

D' L R P M

K.V. A.

/0'4*i0^

Yi^'7l^.:.Q^ \ 7.Q VoU.= .:.7. ^ fi^-S^^B .^/^^pf-d^ ; Arm. A.T. p. pole Max. Fid. AT.

Armaturft. Rev.

o
o

Dia. Outs

Dia. Ins

Gross Length .
Air Vents — L-
Opening Min. (—
Aif- Velocity

Mean

to
to

Xct Length /g-^ x-Sg

Depth b. Slots

Section f$9 Vol.

Flux Density

Loss::fi2.p. cu . C5L. Total
Buried Cu,a5e_Total
Gap Area^7j^-^_: Wts
Vent Area /^^^__:\Vts
Outs. Area ^^^-: Wts

No of Segs
No of Slots

- /„ Mn.Circ.
^X-77=

Section Teeth .
Volume Teeth.
Flux Density.

Loss 'CZ^p. cu £g2^Total

Weight of Iron.

•D

o
o

Mesh Throw

Cond. p Slot

Total Conds ^Smsenes

Jk6_

20 J

0-6

2CLOP0
JQIOO

__ 400
_ ^60

.JSO_

n^JBLJJBjCi^

/org
7o6o

/27 ^
555

jgio _

'3y~50_
L9,0OO
260

J3S.HlQgr^

/'/2

268

Sire of Cond. '23.X^.i2J\j2a s^ cm

Amp. p. sq. C/TL
Length in Slots^
Length outside -35 Sum

Total Length .

Wt. of x.ooo_^4'5_Total
Res. p. 1.000* ^^^Total
Watts p

Surface p.

Watts p. Sq.

^10.

i56'0

^957T
/36

7s Slots :

46 Wound «0

< z - »

\ I

«0

I
I

ic-77>:

7?.. Slots

> <

I *

6950^ 'Sx l'l9x*796'IS7C

Field Stat

Bore

4-6-6

\ Total Air Gap

Gap Co-cff. Kj

Pole Pitch Pole

•5

/• 19

Arc

i6'7

Flux per Pole-
Leakage n.L

f.l

2 '25^/0'

(

\xf^JL3L Flux density

Unbalanced Pull

8750 I

Xo of 9^ 1 /

Mn.Ciic.

1
/54 '

No.ofSloteL-ZiS

x/V =

Vents

66 1

K, . _. .. ..Section

///O '

Weight of Iron

230ki/ofrs.

Shunt. S«pl«s. Coftifti.

Surface p. Watt

1= R

LR.

Amps.

A.T. p Pole n.Load , = , ^___

AT. p. Pnl^fT.nflrl| /w^^. 2600\280O

Surface

No. of Turns t(fUiv,of

Mean 1. Turn .
Total Length .

Resistance

Res. per z. 000.
Size of Cond.

Conds. per Slot-
Total

Length

Wt. per 1,000.

•Total Wt

Watts per Sq..
Star

' 920 920
'r^'95 95

330 I 330'

/06

IPA
50

'0065/ ooas^

7x1-5-

/tf

890

4-5

IM^OL

sp__

890

4-5

Paths in parallel

Magne tizati o n Curve-

Core

Stator Teeth
Rotor Teeth

Gap

Pole Body ^
Yoke

Length

Section

TJpI

mb '4-5

fOio 3^^7_
2760 3

/3/ 1 /5

.7C?.Volts.

A.T.p'»i A.T.

s/g(r_2_ 20

iZ30O. 64j~290

/9.060 35_ '500

69^A

J 970

2S70

•Volts.

A.T.P.

AT.

erriciENCY

Friction and W - _

Iron Loss

Field Loss

Arm. &c. PR

Brush Loss

IJ^Ioad.

Full.

\/-S4.
02

\/ 02

\2 00

'Sf6

Output _

Input

Efficiency

-I

Volts.

A.T.I'. I A.T.

Commutator.

Dia. ^0 Speed l^rpj^sec

Bars _/4^

Volts p. Bar-J

Brs. p. Arm __^

Size of Brs.
Amps p. sq

2^4-5

4.-6

Brush hQ^s72_Q±lSOO.
Watts p. Sq. Q Z5

Mag. Cur. Loss Cur.

Perm. Stat. Slot

., Kot.Slot y =

Zig-zag

X =

X -

X X --^

Amps ; Tot.
: X. =

^'/S, : r. = +

2 X

177
End

Imp. v^ +

Sh. cir. Cur

Starting Torque
Max. Torque _

Max. HP

Slip

Power Factor

618 DYNAMO-ELECTRIC MACHINERY

If we denote the area of the cylindrical working face of the armature by Ag
and the maximum flux-density in the gap by B, then we get the magnetic-loading
equal to AgB. If we have a pole arc equal to 0-7 of the pole pitch, then, as
there are 48 conductors in series and the speed is 12-5 revolutions per second,

70 X 10«=0-7 X 12-5 X 48 X il^B X 0-866.

Observe the multiplier 0*866, which comes into the equation on account of the
circumstance mentioned in paragraph (3) above.

Thus we arrive at the ma^etic loading AgB =0 192 x 10^. If we work the iron
in the teeth at 19,000 lines per sq. cm., we shall require a total mean cross-section
of aU the teeth of 1010 sq. cms. Our conductors, to carry normally 96 amperes
and 25 per cent, over load, may be made 0-23 by 1 -27 cms. Four of these, arranged
as shown in Fig. 533, will require slots about 0-77 x 3-7 cms. To provide room for
72 slots and give the necessary cross-section to the teeth, we shall require a net length
of iron of 16*4 cms. Allowing 11 per cent, for paper on the punchings and 0*6 cm.
for a ventilating duct, we arrive at a gross length of iron of 19 cms. The rest of the
calculation of the armature is the same as for a continuous-current machine, except
in the matter of commutation, which we will consider later. The calculation sheet
is given on page 617. The methods of obtaining the saturations, iron losses, and
cooling conditions are the same as those described on pages 320 and 324. Figs. 529
and 533 give drawings of the machine to scale.

The serieB winding. We must now consider how we are to wind the field poles
so as to give to the excitation its proper phase. The first point to note is that
the six armature circuits are connected in mesh, while the leads from the brush
holders are connected in star.

In Fig. 527 we have a diagram of connections as they would be if the machine
had only three brushes. Obviously this diagram applies equally well to the machine
with six brushes, where brushes at opposite ends of a diameter are in parallel with
one another. The inner circle of Fig. 527 represents the closed winding of the
armature of the advancer. The small letters a, &, c show the three phases mesh
connected. Three brushes — P, Q and R — ^bear on the commutator and convey
the currents to the outer circle, -4, B, C, which represents the winding of the rotor
of the induction motor taken as mesh-connected. It does not matter in practice
whether the rotor of the induction motor is star- or mesh-connected, but for our
diagram it is convenient to connect it in mesh. The arrowheads show the direction
along each conductor, which is taken as positive for the purpose of our clock-diagram
(Fig. 526). P, Q and R are in star, and it is only in series with them that we can
connect the series exciting coils. The voltage in phase A of the rotor is the voltage
we should measure by connecting a voltmeter to the collecting brushes P' and Q'.
In order to make the current in this phase lead, it is necessary to generate a leading
electromotive force in the part a of the armature circuit. From Fig. 526 we foimd
that a suitable e.m.f. to inject into phase A was the e.m.f. XF, which is in phase
with (a -6). From Fig. 527 we see that the current in Q is (6 -a), so that -Q is
(a-h). We will therefore excite the poles under which coils a are passing with
- Q, The span of the armature coils is almost a pole pitch, so that the coils in phase
a will be passing under two adjacent poles, which we will call pole P* and pole

PHASE ADVANCERS

619

•k-

i >

•s

• -

§
1

•

••-

620

DYNAMO-ELECTRIC MACHINERY

Of (see Fig. 530). Now it is not convenient to use only the conductor Q to excite
F" and Q\ because we have to arrange for return paths and also for a compensating
winding, and we want to make a fairly simple mechanical arrangement of the coils.
We therefore take advantage of the known fact that currents

Let us make an arrangement of exciting windings and compensating windings
like that indicated in Fig. 530. There the exciting conductors which pass between
poles F" and Q' are +(?, +(?, -P, -fi. That is to say, they are equivalent to

Fig. 530. — Showing relation of exciting windings and compensating windings to armatuie

windings.

3Q. The question whether the excitation +Q gives a forward or a backward
E.M.F. in a coil depends upon the direction of rotation, and also upon the question
whether the armature is woimd right-handedly or left-handedly. It will be seen
that this arrangement of conductors lends itself to form mechanically a simple
barrel winding as shown in Fig. 531. The conductors lie in two layers, and all the
end connectors of one layer are bent to the right, and all the end connectors of the
other layer are bent to the left. This figure seems fairly complicated, but is made
up by connecting, according to the scheme of Fig. 530, a number of groups of coils
forming part of a simple barrel winding. Fig. 532 shows more exactly how the end
connectors are arranged.

PHASE ADVANCERS 621

Oompetuating wiudiags. The letters in Fig. 530 which ate placed on the salient
poles represent the compensating windings. It is easy to piove that these are in
direct opposition of phase to the
currents in the armature under the
pole. For instance, take the pole
P". The compensating winding on
this is

+ P + P-R-Q, or +3P. ' I

Now the armature coils which ^

he under P* are c and -a, and we |

know that a-c=+P. Moreover, |

the 16 conductors in the pole face s

carrj^ng the currents P, Q and R ^3

are equivalent to 12 conductors J

carrying the P current. Opposite "

the pole P* are 12 armature slots
each carrjdng -2a and 2c. When
we remember that there are two

paths in parallel per phase in the t

armature we see that the currents °

in these 12 slots are exactly balanced
magnetically by the 12P currents
in the compensating winding.

It will be found that an air-gap
of 3 mms. will have an apparent J

length of 3-6 mms. when we take °-

into account the opening of the slots. a

The flux-density in the gap obtained s

by dividing ^^B by Ag is 6950 ; so *

that the ampere-turns on the gap J ■§

will be 1970. The ampere-turns on a ■«

the armature teeth will be 500, on -g

the stator teeth 290, and on the rest g

of the magnetic circuit about 50; ^

so that the ampere-turns per pole « o

will be about 2800 or 5600 per pair • i

of poles. These ampere-turns are .

provided by the 16 conductors which g

thread between the poles P" andQ",
for the 16 conductors carry current
equivalent to 3 x 4P. At its maxi-
mum P is 330 X 1 -41 amperes, which,
multiplied by 12, gives us 5600
ampere-turns per pair of poles. In

practice it will be found unnecessary to adjust the speed exactly, because the
particular power factor at which the motor nms is not a matter of importance.

DYNAMO-ELECTRIC MACHCEEY

lit
el

PHASE ADVANCERS 623

It is not usually necessary to make any provision for the adjustment of the power
factor during running ; it is sufficient that the motor shall take a leading current
from the line at all loads. If it should be necessary to adjust the power factor,
this can be done either by changing the speed of the advancer or by diverting some
of the field current from the series coils. *

Commutation. The most important consideration of the design of the phase
advancer is the obtaining of good commutation. It is chiefly for this purpose that
the field frame and winding described in this paper are provided. Where in a
continuous-current generator the voltage between the bars is small, the commutation
can generally be forced by the resistance of the carbon brushes ; but it is very
much more desirable to provide a commutating e.m.7. which shall at all times be
proportional to the current to be commutated. In the machine here described
this result has been effected by giving each armature coil a span of somewhat less
than the full pitch and arranging the positions of the brushes so that one of the
limbs of each coil is moving in the fringing field of a pole excited by a current
which is at all times proportional to the current under commutation. The currents
in the two branches of the armature, a and - c, which combine to form P, are out
of phase with one another, and are not directly under control of the commutating
flux ; but the rate of change of the current in the coil imder commutation ought
at all times to be proportional to P. Now the pole P' (Fig. 530) is excited so that
the fringing field in which the left-hand limb of the coil a is moving is at all times
proportional to P. By making the coil with a short throw the right-hand limb can
be taken out of the influence of the pole Q'. The exact position for the brushes is,
of course, obtained by trial ; in practice it is found that the commutation is
perfect. The alternation of the current in the armature and field causes a
harmful E.M.7. to be set up in each coil under commutation ; but as the
frequency is so very low (say one cycle per second), this e.m.f. is not sufficiently
great to create any disturbance. In the machine under consideration it only
amounts to one-fourth of a volt.

CHANGE OF SPEED OF INDUCTION MOTORS.

Another use to which these exciters for the rotors of induction motors can be
put is the changing of the speed over a wide range without the necessity for wasteful
rheostats.

In order to change the speed of an induction motor, all that is necessary is to
make the injected e.m.f. XV in Fig. 526 more in phase with the E.M.F. OEa. If
the injected e.m.f. is in the same direction as OEay the speed wiU be increased ;
and if it is opposed to OEa the speed will be decreased. The generation of an in-
jected E.M.F. in phase with OEa is effected by arranging the series coils so that
they carry a component of the current OC, This matter is discussed in the article
referred to below.* At the same time that the speed is increased or diminished,
the power factor can be improved by having a component of the injected e.m.f.
at right angles to OEa»

* Joum. InH, Elec, Engineers, vol. 42, page 599.

624

DYNAMO-ELECTRIC MACHINERY

•—

e
5

c
ea

a
o

PHASE ADVANCERS 625

There are several other systems * of improving the power factor and changing
the speed of induction motors which are of great interest ; but as the matter of
this book has already been extended beyond the limits originally planned by the
publisher, there is not room to consider them here. It is hoped that the author
may have an opportunity of treating in another volume of these and other develop-
ments in the application of dynamo-electric machinery to industrial purposes.

* " Methods of varying the Speed of a.c. Motors,** G. A. Maier, Amer. I.E,E., Proc. 30, p. 2511,
1911 ; " Speed Regulation of 3-Phase Motors,'* G. Meyer, Elekt. Kraflbetr, und Bahnen, 9, pp. 421,
453 and 461, 1911; "Adjustable-speed Polyphase Motors,*' Kn6pfli, Schweiz. eUktrot. Vereir^
Bull. 4, p. 185, 1913.

W. M»

INDEX OF THE CLAUSES IN THE SPECIFICATIONS.

(The first number gives the clause, the second the page.)

A.C. to C.C., converter to run, 248, p. 585.

Acceptance of motor, 188a, p. 469.

Access for repairs, 12, p. 271.

Access to site, difficulty of, 125, p. 461.

Accessibility of power house, 8, p. 271 ; 55,

58, 57, 58 or 59, p. 379.
Accessories not specified,- 178, p. 519.
Air-gap to be stated, 98, p. 441.
Armature, type of, 187, p. 523.
Arrangement, general, of plant, 54, p. 379-

Balance of revolving parts, 85, p. 380 ; 181,

p. 522.
Balancer, converter to act as, 225, p. 563.
Bearing, outboard, 5, ps 271.
Bearings, 87, p. 380 ; 288, p. 590.
Bedplate, 5, p. 271 ; 41, p. 361 ; 70, p. 381 ;

111, p. 443 ; 127, p. 461 ; 186, p. 523 ;

819, p. 611.
Bolts, holding-down, 188, p. 523.
Bonus and penalty on efficiency of rotary

converter, 258, p. 586.
Booster, A.C. See Specification for 1250

K.W. Rotary Converter, p. 560 ; and

see under Converter.
Brushes, 182, p. 501 ; 189, p. 523 ; 285, p. 590.
Brushes, current density in, 288, p. 590.
Brush-gear, 108, p. 442 ; 188, p. 523 ; 191,

p. 524 ; 288, p. 589 ; 806, p. 609 ; 812,

£. 611.
-holders, 18, p. 273 ; 188, p. 501 ; 192,

P- 524-

Cables, 8, p. 271 ; 42, p. 361 ; 78, p. 382 ;

279, p. 591 ; 820, ^.611.
Calibration of measuring instruments, 288,

p. 565.
C.C. side, converter to run well in parallel on,

222, p. 562.
Change in A.C. voltage of converter, 227,

P- 563.
Characteristics of A.C. booster, 215a, p. 561.

of 1250 K.W. rotary converter, 215, p. 560.

of shunt machines, converter to have the,

224, p. 563.

of 750 K.V.A. 3-phase generator, 2, p. 270.

of 2180 K.V.A. 3-pha8e generator, 22, p. 332.

Characteristics — continued,

of 2500 K. V. A. 3-pha8e generator, 82, p. 359.

of 15,000 K.V.A. 3-pha8e turbo-generator,
52, p. 378.

of 2500 K.V. A. 3-phase turbo-generator, 80,
p. 404.

of 75 K.W. C.C. generator, 160, p. 486.

of 1000 K.W. C.C. generator, 155, p. 500.

of C.C. turbo-generator, 178, p. 521.

of 1500 H.P. induction motor, 88, p. 438.

of 350 H.P. induction motor, 120, p. 460.

of 35 H.P. induction motor, 184, p. 4.68.

of 1500 H.P. induction motor witn phase
advancer, 800, p. 608.
Checking of work, 275, p. 591.
Cleaning and painting, 209, p. 528.
Coil, sample, to bo submitted, 270, p. 590.
Coil tested to destruction, 288, p. 335.
Coils, rotor, tests on, induction motor, 97,

p. 440 ; 802, p. 609.
Coils, stator, 91 and 98, p. 439.

stator, insulation, 98, p. 439.
Commutating poles of rotary converter, 248,

P- 585.
Commutation of rotary converter, 229, p. 563 ;

281, p. 589.
of C.C. generator, 184, p. 501 ; 198, p. 5^4«
Commutator, 181, p. 501 ; 188 and 191, p.
523; 282, p. 589; 818, p. 611.
drawing of, to be supplied with tender,

190, p. 523.
grinding gear, 271, p. 59i-
Completion, dates of, 281, p. 392.
Conditions, general, 1, p. 269 ; 21, p. 333 ;

170, p. 519.
Conductors, arrangement of, Ola, p. 439*

insulation of, 289, p. 390.
Connection of generator to engine, 14, p. 270.
Connections of rotary converter, 247, p. 385.
flexible, 285, p. 390.
time for making, 274, p. 391.
and terminals, 287, p. 590.
1250 K.W. Rotary Converter and A.C. Booster,
Specification No. 14 :
Balancer, converter to act as, 225, p. 563.
Booster, characteristics of, 215a, p. 361.
Calibration of instruments, 238, p. 363.
Change in H.T. voltage, 227, p. 363.
Characteristics of booster, 215a, p. 361.
of converter, 215, p. 360.
of shunt machine, 224, p. 363.

628

DYNAMO-ELECTRIC MACHINERY

for, 289,

1260 K.W. Rotary Converter and A.C. Bootter,

Specification No. 14 — continued.
Commutation, 228, p. 563.
C.C. side, running well in parallel on, 222»

p. 562.
Oiverters, voltage drop in, 223, p. 563.
Dutjr of plant, 217, p. 562.
Efl&ciency, 286, p. 564.

calculated, guarantee, 286, p. 565.

measured, guarantee of, 287, p. 565.
Frequency, maintenance of, 220, p. 562,
H.T. voltage, change in, 2S^, p, 563.
Insulation tests, 2I&, p. 564.
Instruments, calibration of, 288, p. 365.

provision of, 289, p. 566.
Interchangeability, 218, p. 562.
Inverted running^^ 219, p. 562.
Leading wattless load, 226, p. 563.
Load, variation of, 221, p. 562.
Noise and vibration, 282, p. 564.
Parallel, running well in, on C.C. side, 222,

p.- 562.
Power-factor control, 216, p. 561.
Power for tests, provision of, 289, p. 566.
Puncture test on site, 284, p. 564.
Running inverted, 219, p. 562.
Shunt machines, characteristic? of, 224,

P- 563.
Stability in operation, 228, p. 563.
Starting, emergency, 280, p. 563.

normal, 281, p. 564.
Tests, instruments and power
p. 566.

insulation, 288, p. 564.

puncture, 284, p. 564.
Variation of load, 221. p. 362.
Vibration and noise, 282, p. 364.
Voltage, H.T., change in, 227, p. 363.

drop in diverters, 228, p. 363.
Wattless load, leading, 226, p. 363.
Work, extent of, 214, p. 360.
2000 K.W. Rotary Converter, Specification
No. 13 :
A.C. to C.C, converter to run, 248. p. 385.
Bearings, 268, p. 390.
Bonus and penalty, 268, p. 386.
Brushes, current density m, 266, p. 390.

type of, 266, p. 390-
Brush gear, 268, p. 389.
Cables, etc., 279, p. 391.
Characteristics, 242, p. 384.
Checking of work, 276, p. 391.
Coil, sample, 270, p. 390.
Commutating poles, 246, p. 383.
Commutation, 261, p. 389.
Commutators, 262, p. 389.
Commutator grinding gear, 271, p. 391.
Completion, dates of, £31, p. 392.
Connections, 247, p. 585.

flexible, type of, 265, p. 390.

and terminals, 267, p. 390.

time for making, 274, p. 391.
Crane, use of, 278, p. 391.
Drawings attached, 282, p. 392.

required, 288, p. 392.
Eflficiency, 249, p. 383.

guarantee, 260, p. 386.

value of I per cent, saved in, 262, p. 386.
Field-regulatmg rheostat, 868, p. 388.

2000 K.W. Rotary Converter, Spscification
No. 13 — continued.
Foundations, 272, p. 391.
Grinding gear for commutator, 271, p. 3Qr.
Hunting, absence of, 260, p. 388.
Insulation, 269, p. 390.
Interchangeability of parts, 276, p. 391.
Load, changing over of, 267, p. 388.
Oscillator, 264, p. 389.
Painting, 278, p. 391.
Penalty and bonus, 268, p. 386.
Poles, commutating, 246, p. 383.
Power-factor variation, 266, p. 387.
Purposes of plant, 241, p. 384.
Rheostat, field-regulating, 268, p. 388.
Samples required with tender, 288, p. 392.
Screw threads, 277, p. 391.
Service, hours per annum. 261, p. 386.
Spare parts, 280, p. 392.
Speed, 246, p. 383.
Starting, 248, p. 383.
Starting motor, 269, p. 388.
Terminals, 867, p. 390.
Tests, 284, p. 392.

notice of, 286, p. 393.
Transformer tappings, wider voltage range

by means of, 2M-266, p. 387.
Voltage range, wider, by means of trana-
former tappings, 264-6, p. 387.
variation of, 266a, p. 387, alteroative
clause.
Work, checking of. 276, p. 391.
extent of, 2M, p. 384.
Corporation, work carried out by, C.C. turbo-
generator, 176, p. 320.
Coupling, 84 and 87, p. 360 ; 110, p. 443.
C.C. tiirbo-generator, 184, p. 323.
half-, 127, p. 461.
Crane, use of, 8, 271 ; 60, p. 379 : 278, p. 591-
Critical speed, 64 n., p. 380; 82, p. 403; 1M»

p. 322.
Current, working, of phase advancer, 811, p.
611.

Delivery of generator, 162, p. 486 ; 218, p. 329.
Diverters, voltage drop in, 228, p. 363.
Drawing of commutator with t«nder, 190,

P- 523.
Drawings attached to specification, 282*

p. 592.

required with tender, 288, p. 392.

Contractor to verify, 174, p. 319.

supplied by Corporation, C.C. turbo*
generator, 210, p. 328.

to be supplied with tender for C.C. turbo-
generator, 211, p. 329.

with specification of induction motor, 116-7,
p. 444.

to be supplied with tender for induction
motor, 119, p. 443.

of site, 172, p. 319.

supplied for Contractoi*s use, 176, p. 319.

where to be seen for purposes of tender, 17ft,

. p. 319.
Drivmg power of phase advancer, 814, p. 611.
Duty of rotary converter, 217, p. 362.

of 13,000 K.V.A. generator, 61, p. 380.

of 1000 K.W. C.C. generator, 167, p. 300.

INDEX OF CLAUSES IN SPECIFICATIONS

629

Duty — continued,

of 34 H.P. indaction motor, 182, p. 462.
of 1500 H.P. induction motor, 86, p. 438.
of motor with phase advancer, 807, p. 610.

Eddy-currents in generator shaft, 68, p. 381.
Efficiency of rotary converter, 285, p. 564 ;

249," p. 585.
of rotary converter, value of i per cent.

saved in, 262, p. 586.
method of determiiung, A.C. generator,

27j, p. 336.
of A.C. generator, 16, p. 272.
of C.C. generator, 166, p. 502 ; 202, p. 525.
guarantee, rotary converter, 250, p. 586.
calculated, guarantee of : rotary converter,

286, p. 565.

measured, guarantee of : rotary converter,

287, p. 565.

of induction motor, guarantee of, 100» p« 441-
of induction motor, method of determining,

99, p. 441 ; 187, p. 469.
of induction motor with phase advancer,

808, p. 609.
E.M.F. wave-form of A.C. generator, 10, p.

271 ; 64, p. 380.
Endurance test of generator, 19h, p. 274 ;

168g, p. 503.
Engine-room in dirty situation, 7, p. 271.
Exciting current measurement, A.C. generator,

27e, p. 335.
Excitation of generator, 11, 271 ; 155i p. 500.
Extent of work, 2, p. 270 ; 81, p. 333 ; 51,

p. 378 ; 80, p. 404.

Factor of safety, 14, p. 272.

Fenders of induction motor, 92, p. 439.

Field-heating run of generator, I9d, p. 273 ;
168c, p. 503.

Field-regufating rheostat for rotary con-
verter, 268, p. 588.

Flywheel for A.C. generator, 24, p. 334.

Foundation plates, 186, p. 523.

Foundations, 6, p. 271 ; 86 and 87, p. 360 ;
74, p. 382 ; 272, p. 591.
of phase advancer, 820, p. 611.

Frame horizontally split, generator, 159, p.
501 ; 179, p. 522.

Framework of generator, 75, p. 382.

Frequency, maintenance of, in rotary con-
verter, 220, p. 562.

750 K.V.A. 8-phase engine-driven A.C. Gen-
erator, Specification No. i :
Access to power-house, 8, p. 271.

for repair, 12, p. 271.
Bedplate and bearings, 5, p. 271.
Brush-holders, 18, p. 273.
Cables, 6, p. 271.
Characteristics, 2, p. 270.
Conditions, general, 1, p. 269.
Connection to engine, 4, p. 270.
Construction, permanent, 14, p. 272.
Crane, use of, 8, p. 271.
Efficiency, 16, p. 272.

750 K.V.A. 8-pha8e engine-driven A.C. Gen-
erator, Specification No. i — continued,
E.M.F. wave form, 10, p. 271.
Endurance test, lOh, p. 274.
Engine-room in dirty situation, 7, p. 271.
Excitation, 11, p. 271.
Field-heating run, 19d, p. 273.
Foundations, 6, p. 271.
Ix>ad, nature of, 8, p. 270.
Magnetization curve test, 19b, p. 273.
Materia], defective, and safety factor, 14,

p. 272.
Oil-tnrowing, 15, p. 272.
Parallel running, 9, p. 271.
Power-house, access to, 8, p. 271.
Puncture tests, 19e, p. 273.
Regulation, 19g, p. 274.
Repair, access for, 12, p. 271.
Resistance tests, 19a, p. 273.
Rheostat, 17, p. 272.
Running conditions, 8, p. 270.
Safety factor of construction, 14, p. 272.
Short circuit, 18, p. 272.
Short-circuit test, 19c, p. 273.
Situation of engine-room, dirty, 7, p. 271.
Slip-rings, 18, p. 273.
Spares, 20, p. 274.
Temperature run, 19f, p. 273.
Tests :

Endurance, 19h, p. 274.

Field-heating run, 19d, p. 273.

Magnetization curve, measurement of,
19b, p. 273.

Puncture tests, 19e, p. 273.

Regulation test, 19g, p. 274.

Resistance test, 19a, p. 273.

Short-circuit test, 19c, p. 273.

Temperature run, 19f, p. 273.
Work, extent of, 2, p. 270.
2180 K.V.A. 8-pha8e gas-engine-driven A.C.
Generator, Specification No. 2 :
Characteristics, 22, p. 332.
Conditions, general, 21, p. 332.
Efficiency, method of determining, 27j,

P- 336.
Exciting current measurement, 27e, p. 333.
Flywheel, 24, p. 334.
Load, nature of, 28, p. 334.
Magnetization curve measurement, 27f, p.

336.
Parallel running, 25, p. 334.
Parallel-running test, 27i, p. 336.
Puncture test, 27d, p. 335.
Shaft, 24a, p. 334.

Short-circuit characteristic test, 27g, p. 336.
Temperature test, 27c, p. 335.
Tests :

Tests after erection, 27a-27j, p. 335.

Tests before shipment, 26, p. 335.

Efficiency test, 27j, p. 336.

Exciting-current test, 27e, p. 335.

Magnetization curve measurement, 27f,

P- 336.
Parallel-running test, 271, p. 336.
Puncture test, 27d, p. 335.
Regulation test, 27h, p. 336.
Short-circuit characteristic test, 27g, p. 336.
Temperature test, 27c, p. 335.
Work, extent of, 21, p. 333.

630

DYNAMO-ELECTRIC MACHINERY

2IM)0 K.V.A. 8-pliase A.C. Generator, SpeciBca-
tion No. 4 :

Bedplate, 41, p. 361.

Cabled, 42, p. 361.

Characteristics, 32, p. 359.

Coupling, 84 and 87, p. 360.

Foundations, 86 and 87, p. 360.

Power house, plan of, 40, p. 360.

Rotor designed for 80 per cent, over speed,
88, p. 360.

Running conditions, 89, p. 360.

Shaft, horizontal, 88 and 86, p. 360.

Star point, 48, p. 361.

Terminals, 44, p. 361.

Work, extent of, 81, p. 359.
16,000 K.V.A. 8-phase Turbo-generator, Speci-
fication No. 5 :

Accessibility, 66» 56, 57, 68 or 60, p. 379.

Arrangement, general, 64, p. 379.

Balance, 65, p. 380.

Bearings, 67, p. 380.

Bedplate, 70, p. 381.

Cables, 78, p. 382.

Characteristics, 62, p. 378.

Crane, use of, 60, p. 379.

Eddy-currents in shaft, 68, p. 381.

Foundations, 74, p. 382.

Framework, 75, p. 382.

Load, nature of, 68, p. 380.

Noise, 78, p. 382.

Purposes of plant, general, 61, p. 380.

Safety, factor of, 66, p. 380.

Shaft, 68 and 60, p. 381.

Site, plan of, 68, p. 378.

Transmission lines, 62, p. 380.

Type of generator, 64, p. 380.

Ventilation, 71, p. 382.

Work, extent of, 61, p. 378.
2600 K.V.A. 8-pha8e A.C. Generator, Specifica-
tion No. 6 :

Characteristics, 80, p. 404.

Critical speed, 82, p. 403.

Load, nature of, 81, p. 405.

Safety factor, 88, p. 405. *

Speed, critical, 82, p. 405.

Work, extent of, 80, p. 404.
76 K.W. Generator, Specification No. 10 :

Characteristics, 160, p. 486.

Delivery, 162, p. 486.

Losses, statement of, 168, p. 486.

Pulley, 161, p. 486.

Tests, 164, p. 487.
1000 K.W. C.C. Generator, Specification No. 11 :

Brushes, 162, p. 501.

Brush-holder, 168, p. 501.

Characteristics, 166, p. 500.

Commutation, 164, p. 501.

Commutator, 161, p. 501.

Duty, 167, p. 500.

Efficiency, 166, p. 502.

Endurance test, 168g, p. 503.

Excitation, 166, p. 500.

Field-heating run, 168c, p. 503.

Frame horizontally split, 150, p. 501.

Iron loss, 168b, p. 503.

Magnetization curve, 168b, p. 502.

Puncture test, 168d, p. 503.

Regulation, 165, p. 301.

Regulation test, 168f, p. 503.

1000 K.W. C.C. Generator, Specification
No. II — continued.
Resistance test, 168a, p. 302.
Rheostat, 167, p. 302.
Short-circuit test, 168b', p. 302.
Spares, 169, p. 303.
Temperature run, 168e, p. 303.
Tests after erection :

Endurance test, 168g, p. 303.

Regulation test, 168f, p. 303.

Temperature run, 168e, p. 303.
Tests at works :

Field -heating nm, 168c, p. 303.

Iron loss, llSb, p. 303.

Magnetization curve, 168b, p. 302.

Puncture test, 168d, p. 303.

Resistance test, 168a, p. 302.

Short-circuit test, 168b', p. 302.
Type of generator, 160, p. 301.
Work, extent of, 168, p. 300.
1000 K.W. C.C. Turbo-generator, Specification
No. 13 :
Accessories not specified, 178, p. 319.
Armature, type of, 187, p. 323.
Balance, 181, p. 322.
Bedplate, 186, p. 323.
Bolts, holding-down, 186, p. 323.
Brushes, 189, p. 323.
Brush gear, 188, p. 323.

fear, construction of, 191, p. 324.
older, sample, 192, p. 324.
Characteristics, 178, p. 321.
Cleaning and painting, 209, p. 328.
Commutation, 198, p. 324.
Commutator, 188, p. 323.

construction of, 191, p. 323.

drawing of, with tender, 190, p. 323.
Conditions, general, 170, p. 319.
Corporation, work carried out by, 176, p. 520.
Coupling, 184, p. 323.
Critical speed, 180, p. 322.
Delivery, 218, p. 529.
Drawing of commutator with tender, 190,

P- 523-
Drawings, alternative clause, 176, p. 319.

Bupphed by Corporation, 210, p. 328.

to be supplied with tender, 211, p. 329.
Efficiency, 202, p. 323.
Extent of sections of specification, 177, p.

320.
Frame horizontally split, 179, p. 322.
Holding-down bolts, 186, p. 323.
Instruments, provision of, 208, p. 328.
lioad, throwing on and off, 194, p. 324.
Load and steam, provision of, 207, p. 328.
Machines, similar, in operation, 196, p. 324
Maintenance period, 206, p. 328.
Noise, 182, p. 322.
Painting and cleaning, 209, p. 328.
Parallel running, 197, p. 323.
Provisional sum, 212, p. 329.
Rating, 178, p. 321.
Safety factor, 188, p. 322
Spares, 200, p 323.

Specification, extent of sections, 177, p 320.
Speed, critical, 180, p. 322.
Steam and load, provision of, 207, p. 328.
Temperature rise, measurement of, 205, p.
527.

INDEX OF CLAUSES IN SPECIFICATIONS

631

1000 K.W. C.C. Tarbo-«enerator, Specification
No. 13 — continued.
Terminals, 199, p. 525.
Tests after delivery, 204, p. 527.
at makers' works, 208, p. 526.
Tools, 201, p. 525.
Ventilation, 196» p. 524.
Work carried out by Corporation, 178,
p. 520.
extent of, 171, p. 519.
Grinding gear for commutator, 271, p. 591.

H.T. voltage of converter, change in, 227,

P- 563.
Holding-down •bolts, C.C. turbo-generator,

186, p. 523.
Hunting, absence of, in rotary converter, 260,

p. 588

Induction motor, see " Motor."
Information, general, phase -advancer specifi-
cation, 815, p. 611.
Instruments, calibration of, 288, p. 565.

for tests, provision of, 181, p. 462 ; 208,
p. 528 ; 289, p. 566.
Insulation of conductors, 269, p. 590.
of stator coils, 98, p. 439.
tests on rotary converter, 288, p. 564-
Interchangeability of rotary converters, 218,
p. 562.
of parts, 276, p. 59i-
Inverted rotary converter, 219, p. 562*
Iron-loss test, C.C. generator, 16i8b, p. 503.

Leading wattless load of rotary converter,

226, p. 563.
Load of rotary converter, changing over of,
257, p. 588.
of generator, nature of, 8, p. 270 ; 28,

P- 334 ; «8, p. 380 ; 81, p. 405.
of induction motor, nature of, 121, p. 460.
and steam for testing, provision of, 207,

p 528.
throwing on and off, C.C. turbo-generator,

194, p. 524.
variation of : rotary converter, 221, p. 562.
Losses in C.C. generator, 158, p. 486.
in phase advancer, 822, p. 612.

Machines, similar, in operation, C.C. turbo-
generator, 195, p. 524.
Magnetization curve of generator, 19b, p. 273 ;

27f, p. 336 ; 168b, p. 502.
Maintenance period, 188b, p. 469 ; 206, p. 528.
Material, defective, and safety factor : A.C.

generator, 14, p. 272.
Measurements, Contractor to make, 118, p. 444.

on site by Contractor, 174, p. 519.
Mine, carriage of motor through, 128, p. 461.

plan of, 126, p. 461.
1500 H.P. 8-pha8e 1850 K.V.A. Induction
Motor, Specification No. 7 :
Air-gap, 98, p. 441.
Bedplate, 111, p. 443.

1500 H.P. 8- phase 1850 K.V.A. Induction
Motor, Specification No. 7 — continued.
Brush gear, 106, p. 442.
Characteristics, 88, p. 438.
Coils, stator, 91 and 98, p. 439.

stator, insulation of, 98, p. 439.

rotor, tests on, 97, p. 440.
Conductors, arrangement of, 91a, p. 439.
Coupling, 110, p. 443.
Drawings with specification, 116-7, p- 444.
Drawings to be supplied with tender, 119,

. P- 445-
Efficiency, 99, p. 441.

guarantee of, 100, p. 441.
Fenders, 92, p. 439.
Function of motor, 86, p. 438.
Guarantee of efficiency, 100, p. 441*
Insulation of stator coils, 98, p. 439.
Measurements, Contractor to make, 118,

p. 444.
Motor, function of, 86, p. 438.
Pressure tests, 94, p. 440.
Power factor at starting, 104, p. 442.
Rotor, 96, p. 440.

coils, tests on, 97, p. 440.

type of, 87, p. 438.
Short-circuiting device, 107, p. 442.
Slip-rings, 105, p. 442.
Spares, 114, p. 444.
Starting, 108, p. 442.
Stator frame, 90, p. 439.

coils, 91 and 98, p. 439.

coils, insulation of, 98, p. 439.

winding, tests on, 95, p. 440.
Starting, 108, p. 442.

power factor at, 104, p. 442.
Terminals, 112 or 118, p. 443.
Tests, pressure, 94, p. 440.

on stator winding, 95, p. 440.

on rotor coils, 97, p. 440.

at makers* works, 101, p. 441.

on site, 102, p. 441.
Tools, 115, p. 444.
Ventilation, 108, p. 443.
850 H.P. 8-pha8e 805 K.V.A. Induction Motor,
Specification No. 8 :
Access, difficulty of, 125, p. 461.
Bedplate, 127, p. 461.
Carriage through mine, 128, p. 461.
Characteristics, 120, p. 460.
Coupling, half-, 127, p. 461.
Instruments for tests, 181, p. 462.
Load, nature of, 121, p. 460.
Mine, plan of, 126, p. 461.
Power-factor test, 180, p. 462.
Rheostat, separate quotation for, 123,

p. 461.
Situation, 124, p. 461.
Speed, variation of, 122, p. 460.
Test of power factor, 180, p. 462.
Tests on site, 129, p. 461.

instruments for, 181, p. 462.
85 H.P. 8-pha8e Induction Motor, Specifica-
tion No. 9 :
Acceptance, 188a, p. 469.
Characteristics, 184, p. 468.
Efficiency, 187, p. 469.
Maintenance period, 188b, p. 460.
Pulley and slide-rails, 186, p. 468.

632

DYNAMO-ELECTRIC MACHINERY

— o

85 HJ. 8-phase Induction Motor, Specifica-
tion No. 9 — continued.
Purpose of motor, 182, p. 468.
Rotor, type of, 188, p. 468.
Slide-rails and pulley, 185, p. 468.
Tests, 188, p. 469.
Work, extent of, 186, p. 468.
Motor, induction, with phase advancer. See
under '* Phase Advancer " and Specifica*
tion 7a,
duty of, 807, p. 610.
desired power factor of, 808a, p. 610.
power factor of, 808, p. 610.
rating of, 806, p. 610.

Noise of machines, 72, p. 382 ; 182, p. 522 ;
282, p. 564.

Oii-thro¥ang, 15, p. 272 ; 828, p. 612.
Oscillator for rotary converter, 264, p. 589.

Painting and cleaning, 209, p. 328 ; 278,

p. 591.
Parallel, converter running well in, on C.C
side, 222, p. 362.
running of generator, 0, p. 271 ; 25, p. 334 ;

197, p. 523-
running test, A.C. generator, 271, p. 336.
Penalty and bonus for efficiency, 258, p. 386.
1500 H.P. 8-pha86 1850 K.V.A. Induction Motor
with Phase Advancer, Specification
No. ja:
Brush gear, 805, p. 609.
Characteristics, 800, p. 608.
.^ Coils, rotor, tests on, 802, p. 609.
Efficiency, 808, p. 609.
Tests on rotor coils, 802, p. 609.

on site, 804, p. 609.
Work, extent of, 801, p. 609.
29 K.V.A. 8-phase 50-volt Phase Advancer,
Specification No. 16 :
Bedplate and bearings, 819, p. 611.
Brush gear and commutator, 816, p. 611.

and slip-rings, 812, p. 611.
enables, 820, p. 611.
Commutator, 816, p. 611.
Current, working, 311, p. 6ti.
Driving power of phase advancer, 814,

p. 611.
Duty of motor, 807, p. 610,
Foundations, 820, p. 611.
Information, general, 815, p. 611.
Losses in phase advancer, 822, p. 612.
Motor, duty of, 807, p. 610.

Sower factor of, 808, p. 610.
esired power factor, 808a, p. 610.

rating of, 806, p. 610.
Oil-throwing, 828, p. 612.
Power factor of motor, 808, p. 610.

desired, 808a, p. 610.
Puncture test, 818, p. 611.
Slip at full load, 810, p. 610.
Slip-rings, 812, p. 611.
Temperature rise, 817, p. 611.
Test, puncture, 818, p. 611.

29 K.V.A. 8-phase 50-volt Phase Advancer,

Specification No. 16 — continued.
Voltage, standstill, 809, p. 610.
Work, setting machine to, 821, p. 611.
Poles, commutating, of rotary converter, 246,

P- 585-
Power for tests, provision of, 289, p. 366.

Power factor control : rotary converter, 216,

p. 561.
Power factor of motor with phase advancer,
808, p. 610.
desired, of motor with phase advancer,

808a, p. 610.
at starting 1300 H.P. induction motor, 104,

p. 442.
test of induction motor, 180, p. 462.
variation in rotary converter, 256, p. 387.
Power house, access to : A.C. generator, 8,
p. 271.
plan of, 2300 K.V.A. generator, 40, p. 360.
Pressure tests of induction motor, 94, p. 440.
Provisional sum, 212, p. 329.
Pulley for generator, 151, p. 486.

and slide-rails for motor, 185, p. 468.
Puncture test, 19e, p. 273 ; 26a, p. 333 ;
27d, p. 333 ; 168d, p. 303 ; 616, p. 611.
on site, 284, p. 364.

Regulation of C.C. generator, 166, p. 301.

test, 19g, p. 274 ; 168f, p. 503.
Repair of generator, access for, 12, p. 271.
Resistance tests on generator, 19a, p. 273 ;

168a, p. 302.
Rheostat, neld-regulating, for rotary converter,
258, p. 388.
for A.C. generator, 17, p. 272.
for C.C. generator, 167, p. 502.
for induction motor, separate quotation for,
128, p. 461.
Rotary converters. See under " Converter.*'
Rotor of induction motor, 87, p. 438 ; 96,
p. 440 ; 188, p. 468.
designed for 80 per cent, over speed, 2500

K.V.A. generator, 89, p. 360.
coils of induction motor, tests on, 97,
p. 440 ; 802, p. 609
Running conditions of generator, 8, p. 270 ;
89. p. 360.
inverted : rotary converter, 219, p. 362.

Safety factor of construction, 14, p. 272 ; 66,
p. 380 ; 88, p. 405 ; 188, p. 322.

Samples required with tender for rotary
converter, 288, p. 392.

Screw threads, 277, p. 39i«

Service, hours per annum, of rotary con-
verter, 251, p. 386.

Shaft of generator, 24a, p. 334 ; 88 and 85,
p. 360 ; 68 and 69. p. 381.

Short-circuit, A.C. generator, 18, p. 272.
characteristic test of A.C'. generator, 27ff,

P- 336-
test of A,C. generator, 19c, p. 273.

teat on C.C. generator, 168b', p. 502.

test, 19c, p. 273 ; 168b', p. 302.

Short -cireuiting device on induction motor,

107, p. 442.

INDEX OF CLAUSES IN SPECIFICATIONS

633

Shunt machines, converter to have . charac-
teristics of, 224, p. 563.
Site, drawings of, 172, p. 519.

measurements on, by Contractor, 174, p.

519.
plan of, 68, p. 378.
Situation of engine-room, dirty, 7, p. 271.

of machine, 124, p. 461.
Slide-rails and pulley for induction motor,

186, p. 468.
Slip at full load, phase advancer, 810, p. 610.
—^lip-rings, 18, p. 273 ; 106, p. 442 ; 812,
p. 611.
Spare parts, 20, p. 274 ; 114, p. 444 ; 169,

p. 505 ; 200, p. 525 ; 280, p. 592.
Specification, general conditions of, 1, p. 269 ;

21, p. 333 ; 170, p. 519-
Speed of rotary converter, 246, p. 585.

critical, 82, p. 403 ; 180, p. 522.
Stability in operation of converter, 228, p. 563.
Star point, 2500 K.V.A. generator, 48, p. 361.
Starting of rotary converter, 281, p. 564 ;
248, p. 585.
emergency, of rotary converter, 280, p. 563.
1500 H.P. induction motor, 108, p. 442.
power factor at, 1500 H.P. induction motor,

104, p. 442.
motor for rotary converter, 269, p. 588.
Stator coils, insulation of, 98, p. 439.

rigidity of, 91, p. 439.
Stator frame of 1500 H.P. induction motor,
90, p. 430.
winding, tests on, 96, p. 440.
Steam and load, provision of, 207, p. 328.

Temperature rise, measurement of, C.C. turbo-
generator, 206, p. 327.
rise, phase advancer, 817, p. 611.
run of generator, 19f, p. 273 ; 27c, p. 335 ;
168e, p. 303-
Terminals, 44, p. 361 ; 112 or 118, p. 443 ;

199, p. 323 ; 267, p. 390.
Test of coil to destruction, 26b, p. 333.
of insulation, 288, p. 364.
puncture, 818( p. 611.
on site, ^4, p. 364.
Tests on converter, 284, p. 392.

on generator, after erection, 27a-27), p 333 ;

168e to g, p. 303 ; 204, p. 327.
on generator, 10, p. 273 ; 164, p. 487.
instruments and power for, 289, p. 366.
at makers* work?, C.C. generator, 168a to d,
p. 302 ; 208, p. 326.

Tests at makers* works of 1300 H.P induction
motor, 101, p- 441-
on 34 H.P. induction motor, 188, p. 469.
notice of, 286, p. 393-
pressure, voltage of, 94, p. 440.
on rotor coils, 97, p. 440 ; 802, p. 609.
before shipment of A.C. generator, 26, p. 333.
on site of induction motor, 102, p. 441 ;

804, p. 609.
on stator winding of 1300 H.P. induction
motor, 96, p. 440.
Tools, 116, 444 ; 201, p. 523.
Transformer tappings, wider voltage range of

converter by means of, 264-6, p. 387.
Transmission lines, 62, p. 380.
Turbo-generator. See under "Generator."

Variation of load of rotary converter, 221,

p. 362.
Ventilation, 71, p. 382 ; 108, p. 443 ; 196,

p. 524.
Vibration and noise, 282, p. 364.
Voltage drop in diverters : rotary converter,

228, p. 363.
H.T., change in : rotary converter, 227,

p. 563-
range, wider, of rotary converter, by means

of transformer tappings, 264-6, p. 387.
standstill, of phase advancer, 809, p. 610.
variation of, in rotary converter, 266a,

P- 587 (alternative clause).

Wattless load, leading, of converter, 226,

P- 563.
Wave-form of E.M.F., 10, p. 271 ; 64, p. 380.
Work carried out by Corporation, C.C. turbo-
generator, 176, p. 320.
checking of, 275, p. 391.
extent of, rotary converter, 214, p. 360 ;

240, p. 384.
extent of, A.C, generator, 2, p. 270 ; 22,

p. 333; 81, p. 359; 61, p. 378; 80,

p. 404.
extent of, C.C. generator, 168, p. 300.
extent of, C.C. turbo-generator, 171, p. 319.
extent of, 34 H.P. induction motor, 186,

p. 469.
extent of, induction motor with phase

advancer, 801, p. 609.
setting phase advancer to, 821, p. 611.

O —

GENERAL INDEX

Abrasion of insulation, risk of, 194

Air. See also Ventilation

Air, baffled, heat conductivity of, 220

cooling by, 229

di-awn in at ends and expelled radially, 206

relation between volume and weight of, 244

required for cooling, 206, 390

temperature rise of, 249, 394

throttled, effect of, 247

velocities in various parts, 242, 391

watts required to heat, 216, 247
Air and iron, 71
Air-gap, the, 56, 62 (and see Calculation sheets)

ampere-turns on, 63, 417 (and see Calcula-
tion sheets)

contraction coefficient K,j^ 66, 417

dimensions, 326

flux-density in, 56, 65, 77, 281, 464 (and see
Calculation sheets)

of induction motor, 472

length of, 62, 347, 416, 449

length of : effect on zigzag leakage, 424

between punchings, 84

of rotary converter, 574
Air-gap-and -tooth-saturation curve, 76, 377,

395
Air pockets, presence of causing formation of

nitric acid. 192
Air-space, effect of, 71

effect on tooth reluctance, 71
AUen^ V. M., on design of moulds for coils,

154 to 168 (see Preface)
AUgemeine Elektricitdts OeseUschaft, 369, 406
Alternator. See Generator
Alternator, synchronous, running in parallel
with network, 337

wave-form of, 27
Aluminium, 135
American Institution of Electrical Engineers,

testing rules, 189
Amortisseur, 411

design of, 35% 578, 652

on rotary converter, 602
Amperes per brush arm, 488, 505, 530

per conductor, 512

per slip-rin^ and amperes per terminal, ratio
between m converters, 541
Ampere -turns, 36

absorbed in pole -body, 328

on air-gap, 62, 330, 417 (and see Calculation
sheets)

on air-gap and teeth, 77

Ampere-turns — continued

of armature, 279, 282, 286, 478, 518 (and see

Calculation sheets)
of armature, relation to no-load field A-T.,

290, 293
to drive full-load armature current on short-

circuit, 283
required for break-joint, 83
added to excitation of converter in order to

obtain a given rise in voltage, 597
on core, 82, 330, 419
increase due to leakage on load, 281, 331,

358, 398
incre€ksed, due to extra leakage on load, 281,

358, 398
mean value of in S-phase armature, 280
no-load field : relation to armature A.T.,

290,293
on pole, what dependent upon, 299
for pole body, 330
per pole, 327, 493, 508, 599
per pole, effect of diam. of A.C. generator,

302
per pole, effect on, of speed, 303
per pole, effect on, of widening frame, 234,

303
of shunt coil, 141
on teeth, 73, 330, 395, 418
on yoke, 85
Ampere-wires per cm., 383, 532

per inch of perimeter, 383, 532
Angermann on eddy-currents, 145 n.
Angle made by coil-end with iron, 157

of displacement of synchronous machines,

339, 342
of lag in induction motor, 414
between neutral plan&s, 349
between centre Ime of pole and phase line

of terminal voltage, 342
Angular irregularity of engine, 339, 603
Annealing sheet iron after punching, 53
Apparent fiux-densities, curve showing, 72
Armature of A.C. ^nerator, 274
cm rent density m, 490
demagnetizing effect of, 279
resistance of, 454
ring, 516

ampere-turns, 279, 282, 286, 478, 518
circuits in parallel : number of =2a, 512
Armature coils, design of, 151
insulation of, 199
short type : mould for, 153
specification of mould for, 161, 167

GENERAL INDEX

635

Armature coils — continued

of strap, 162, 480

of wire, 152
Armature core, 82
Armature windings, 87

3 -phase, classes of, 101

of C.C. generators, 475

of 10,000 K.V.A. 3-phase generator, 119
Amoldy E,y on heating of coils, 230 n.

singly re-entrant multiplex winding, 511, 515
Arnold and La Cour on commutation, 480 n.
Asbestos, qualities of, 176

solid-filled, permissible temp., 256
Asbestos-slate, qualities of, 176
Asphalt um compounds little afifected by pro-
ducts of discharge, 192

Baily, F. C, on hysteresis, 45 n.
Bakellte, qualities of, 176, 178
Balancing flux, 452

rings, 513
Barling, W, H., on magnetizing current, 280 n.
Barlow, T. M., on heat conductivity, 253 n.
Barr, J. R., on parallel running of alternators,

337 n.
Barr and Archibald on A.C. machinery, 542 n.
Barrel-winding, 116, 427, 620

clamping of, 134

throw of, as affected by number of poles, 1 18
Barrel end-connectors, 115
'* Base " circle, 164, 166
De Bast, 0., on interpoles, 480 n.
Bauer, R., on power-mctor improvement, 612 n.
Beaiiie, Dr. R., on harmonic analysis, 22
Beekman, R. A., on heating, 255 n.
Benischke, G., on parallel running of alter-
nators, 337 n.

on induction motor, 421 n.
" Bent ends " of connectors, 94
Bevel on mould, 161
Bitumen, 178

Blondel, ^., on synchronizing, 337 n.
Booster, A.C., 546, 695

connections of, 647

design of, 679

driving of, 679

E.M.F. of, 581

frame, size of, 580

position of, 646

reaction of, 679

field winding of, 683
Boucheroty T., on irregularities in speed, 337 n.
Boulardet, E., on heating of electrical ma-
chinery, 266 n.
Boulding, R, 8. U., on ripples in wave -form,

306 n.
Bragstad and Frdnkel on iron losses, 86 n.
Brass, heat conductivity of, 220
Breadth coefficient, 112, 307
Breakdown due to disruption of molecules, 179

due to heating due to electric conduction, 179

on insulation, experience of, 193

over surface of insulation, 196

te8t«, 187

between turns, 197

due to high voltage, 196

Break-jointfi, 83

uneven, 50
" Breathing " action of insulated coil, 191
Breslauer, m., on iron, 86 n.
British Electrical and Allied Manufacturers*

Association, 188
British Thomson- U&iiston Company, Ltd., 119
Brow leakage, 426
Brown, Boveri <«? Co., 3-phase motor, 209

scheme of ventilation, 207
Brunt, ^., on design of auxiliary poles, 480 n.
Brush arms : distance between, 517, 530
Brush discharge, 192

effect of, on insulation, 186
Brush -^ar, 482, 576
Brush-holders, 482

" Box type," 482
Brushes of carbon, 482, 516

and commutator, contact between, 482, 516
Brushes, contact area, 578

of copper-carbon, 484

graphitic, 484

metal, 482

pit<;h of, for converters, 539

resistance of, 477, 483

rocked too far back, 494

voltage drop at, 578

width of, 476, 479, 499

width of, for converter, 575
'* Buried copper " losses, 324
Buyers and efficiency, 495

Calculation sheets, see list of, p. xiii

generator, 316
Calculation sheet, general form of, 317
Calculations, preliminary : saving of time in,

293
Calico, safe mechanical pressure on, 195
Caminati, C, on heating of electrical ma-

chinerv, 265 n.
Campbell, ^., on iron, 86 n.
Campos, 0., on phase variation, 612 n.
Canvas, oiled : qualities of, 176
Capacity, specific inductive, of insulating

materials, 186
Carbon, effect of, in steel, 44
Carbon brushes. See Brushes
Le Carbone Company, 483
Carter, F. W., on fringing of flux, 14 n.,

64 n.
Castings, delivery of, 40

C.C. to A.C., running : rotary converter, 568
Cell of insulation, 164, 160
Cellulose insulation, 183, 189, 190
Centre line of flux, 280
Centrifugal forces on water- turbine -driven

generator, 361
on steam-turbine-driven generator, 366, 529
Change-over switches on transformer, 649
Chorded winding, eddy current in, 119
Chording the winding, effect of. 111
Christie, J., on ventilation, 204 n.
Chubb, L. W., on heating of electrical machin-
ery, 255 n.
arcle, " base," 164, 166
Circle diagram of induction motor, 413, 468

636

DYNAMO-ELECTRIC MACHINERY

Clock diagram, 456

of A.C. generators in parallel, 338

of rotary converter, 695, 598

of leading current in rotor of induction
motor, 613, 614
Coils, armature : cooling of, 389

breadth of, 306

circular, 495

concentric, 151

concentric, formers for, 168

cylindrical, of cotton -covered wire, cooling
of, 231

of diamond shape, 151

diamond- type : design of, 156, 157

end of, cooling coefficient hg, 232

external insulation of, 200

insulation of, 199

involute, 151

lattice type, 151

length of mean turn of, 143

mechanical protection of, 194

moulds or formers for, 151, 152

mush-wound, insulation of, 201

number of, in C.C. generator, 475

number of turns in, 511

average overhang of, 426

fitting tightly on poles, 231

projection of, 171

" puUed," 152

rectangular, 495

room in slot for external wrapping, 202

" sections " of, 155

short throw, 481

short-type, design of, 163

short-type, calculation sheet for, 168

sides of, cooling coefficient hi, 233

skew, 96

stator : insulation of, 202

of strap, 152

of strap : mould for, 162

taping of, 200

temperature distribution in, 226

wire-wound, 480
Coil-end, angle made with iron, 157
Coil-pitch, calculation of, 159
C'Oil-surface, watts dissipated from, 331, 234
Coil-throw, 156
Commutatine pole, 498, 510, 518, 530

axial length of, 480

leakage flux, 479

windmg of converter, 574
Commutating windings, 536
Commutation, 475, 476

of a two-turn coil, 511

curves of, 477
Commutation affected by phase- swinscing, 600,

603
Commutation of phase advancers, 623
Commutator, 498, 510

of converter, 575

cooling surface of, 499

diameter of, 530

eccentricity of, 482

flat in, 482

length of, 594

radial type, 516, 536

speed, 540

throw on, =y, 512

wear on, 482

Commutator bars, method of making connee*
tions to, 537

intermediate, 517

number of, =Kfn, 511, 512, 517, 530
Commutator bars : doubling the number of,
517

number of, per pole, 511

per pole on converter, 569

voltage on, 517

critical voltage of, 532

width of, 479
Commutator and brushes, contact between,

482, 516
Commutator-necks, 537
Compensating windings, 518, 536

of phase advancer, 621
Compounding of rotary converter, 550
Compound-wound generators, 484
Concentric coils, 121

formers for, 168
Concentric connections, 88
Concentric winding, 427
Conductor diagram, 87, 92

diagram for 3-pha8e winding, 93

size of, in converter, 571
Conductors, armature, of aluminium, 136

heating of, 541

reducing number of, 513

of rotary converter, 545

arrangement of, 193

arrangement of, in armature slots, 138

arrangement of, on field-magnets, 139

depth of, in slots : effect on eddy-current,
149

double, to reduce eddy -current loss, 390

heat conduction along, 225

insulation and assembly of, 198

laminated, 150

material for, 134

number of, on A.C. generator, 320

number of, on C.C. generators, 511

number of, on induction motor, 449

in parallel, 25

per pole, odd number of, 581

of rectangular cross-section, 490

in rotor of induction motor, 466

shape of, 138

size of, 141, 149, 389, 513, 532

space occupied by, 137

stranded, 150

twisted, 393
Conductors of armature : size of, 349, 508 (and

see Calculation sheets)
Conductivity of insulation, 179
Connections to back of armature, 517, 537

concentric, on 3-pha8e winding in 3 tiers, 93

Walker's auxiliary, on C.C. turbo-gpnerator,
517
Connector diagram, 87

Connectors, auxiliary, to intermediate com-
mutator bars, 517
Connectors, end-, barrel, 371

involute, 537

stranded copper, 390
Construction, type of : purchaser's preference

for, 262
Continuous-current generator, design of, 487
Continuous-current generators, 475 (see page

zni)

GENERAL INDEX

637

Contraction coefficient Kg, 65, 78
Contraction ratio, 64

Controller for varying voltage to rotary, 549
Converters, rotary, 539

to run C.C. to A.C., 579

connections, 554-558

2000-K.W., for electrolytic work, design of,
594

efficiency of, 578

four-phase : ampere ratio, 541

foar-phase : voltage ratio, 540

frequency, regulation of, 579

rotary, instead of C.C. turbo-generator, 516

hunting of, 600

over-compounded, 596

over-excited, 596

parallel running, 600

single-phase : ampere ratio, 541

single-phase : voltage ratio, 540

six-phase : ampere ratio, 541

six-phase : voltage ratio, 540

small, 604

starting of, 604

when suitable, 539

three-phase : ampere ratio, 541

three-phase : voltage ratio, 540

variation of voltage by variation of excita-
tion, 595

under-excited, 596
Cooling, air required for, 216, 390

air, velocity of, 21 1

of armature coils, 389

by air, 229

coefficient of armature, 229
of field-ooil, 229, 232, 233
of cylindrical field-magnet, 229
of iron surface of duct, 229
Arf,230
of coil-ends, 232

conditions, rules for predetermining, 254

of field-coils, 349

of induction motor, 472

as affecting breakdown of insulation, 179

of cylindrical coil, 231

of brass cylinder, 231

of rotating field-coils, 232

of stator, 324, 392

of external surface of stator, 254

surface, 388

surface : copper and aluminium compared,
137

surface of shunt coil, 142

root of teeth, 210

of turbo-generator, 225, 394

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